Nuscale Power, LLC Submittal of the Approved Version of Nuscale Topical Report, “Loss-Of-Coolant Accident Evaluation Model,” TR-0516-49422, Revision 2

LO-0620-70792

July 7, 2020 Docket No. PROJ0769

U.S. Nuclear Regulatory Commission ATTN: Document Control Desk One White Flint North 11555 Rockville Pike Rockville, MD 20852-2738

SUBJECT: NuScale Power, LLC Submittal of the Approved Version of NuScale Topical Report, “Loss-of-Coolant Accident Evaluation Model,” TR-0516-49422, Revision 2

REFERENCES: 1. NRC Letter to NuScale, “Final Safety Evaluation for NuScale Power, LLC Topical Report TR-0516-49422, Revision 2, ‘Loss-of-Coolant Analysis Model,’” dated June 22, 2020 (ML20181A270)

2. Letter from NuScale to NRC, “NuScale Power, LLC ‘Submittal of Errata to Loss-of-Coolant Accident Evaluation Model,’ TR-0516- 49422, Revision 2,” dated June 19, 2020 (ML20175A345)

3. Letter from NuScale to NRC, “NuScale Power, LLC Submittal of ‘Loss- of-Coolant Accident Evaluation Model,’ TR-0516-49422, Revision 2,” dated May 27, 2020 (ML20148T471)

By referenced letter dated June 22, 2020, the NRC issued a final safety evaluation report documenting the NRC Staff conclusion that the NuScale topical report “Loss-of-Coolant Accident Evaluation Model,” TR-0516-49422, Revision 2, is acceptable for referencing in licensing applications for the NuScale small modular reactor design. The referenced NRC letter requested that NuScale publish the approved version of TR-0516-49422 within thirty days of receipt of the letter.

Accordingly, Enclosure 1 to this letter provides the approved version of TR-0516-49422-P-A, Revision 2. The enclosure includes the June 22, 2020 NRC letter and its final safety evaluation report.

Enclosure 1 contains proprietary information. NuScale requests that the proprietary version be withheld from public disclosure in accordance with the requirements of 10 CFR § 2.390. The enclosed affidavit (Enclosure 3) supports this request. Enclosure 1 has also been determined to contain Export Controlled Information. This information must be protected from disclosure per the requirements of 10 CFR § 810. Enclosure 2 contains the nonproprietary version of the approved topical report package.

This letter makes no regulatory commitments and no revisions to any existing regulatory commitments.

NuScale Power, LLC 1100 NE Circle Blvd., Suite 200 Corvallis, Oregon 97330 Office 541.360-0500 Fax 541.207.3928 www.nuscalepower.com LO-0620-70792 Page 2 of 2 07/07/2020

If you have any questions, please contact John Fields at 541-452-7425 or at [email protected].

Sincerely,

Zackary W. Rad Director, Regulatory Affairs NuScale Power, LLC

Distribution: Gregory Cranston, NRC Prosanta Chowdhury, NRC Michael Dudek, NRC Rani Franovich, NRC

Enclosure 1: “Loss-of-Coolant Accident Evaluation Model,” TR-0516-49422-P-A, Revision 2, proprietary version Enclosure 2: “Loss-of-Coolant Accident Evaluation Model,” TR-0516-49422-NP-A, Revision 2, nonproprietary version Enclosure 3: Affidavit of Zackary W. Rad, AF-0620-70793

NuScale Power, LLC 1100 NE Circle Blvd., Suite 200 Corvallis, Oregon 97330 Office 541.360-0500 Fax 541.207.3928 www.nuscalepower.com LO-0620-70792

Enclosure 1:

“Loss-of-Coolant Accident Evaluation Model,” TR-0516-49422-P-A, Revision 2, proprietary version

NuScale Power, LLC 1100 NE Circle Blvd., Suite 200 Corvallis, Oregon 97330 Office 541.360-0500 Fax 541.207.3928 www.nuscalepower.com LO-0620-70792

Enclosure 2:

“Loss-of-Coolant Accident Evaluation Model,” TR-0516-49422-NP-A, Revision 2, nonproprietary version

NuScale Power, LLC 1100 NE Circle Blvd., Suite 200 Corvallis, Oregon 97330 Office 541.360-0500 Fax 541.207.3928 www.nuscalepower.com Loss-of-Coolant Accident Evaluation Model

TR-0516-49422-NP-A Rev 2 Contents

Section Description

A Letter from NRC to NuScale, “Final Safety Evaluation for NuScale Power, LLC Topical Report TR-0516-49422, Revision 2, ‘Loss-of-Coolant Analysis Model,’” June 22, 2020 (ML20181A270)

B NuScale Topical Report: Loss-of-Coolant Accident Evaluation Model, TR-0516-49422-NP-A, Revision 2

C Letters from NuScale to the NRC, Responses to Requests for Additional Information on the NuScale Topical Report, “Loss-of-Coolant Accident Evaluation Model,” TR-0516-49422

D Letter from NuScale to NRC, “NuScale Power, LLC Submittal of ‘Loss-of-Coolant Accident Evaluation Model,’ TR-0516-49422, Revision 2,” dated May 27, 2020 (ML20148T471)

TR-0516-49422-NP-A Rev 2

Section A

Mr. Zackary W. Rad Director, Regulatory Affairs NuScale Power, LLC. 1100 Circle Boulevard, Suite 200 Corvallis, OR 97330

SUBJECT: FINAL SAFETY EVALUATION FOR NUSCALE POWER, LLC TOPICAL REPORT TR-0516-49422, REVISION 2, “LOSS-OF-COOLANT ANALYSIS MODEL”

Dear Mr. Rad:

By letter dated December 30, 2016, the applicant, NuScale Power, LLC (NuScale) submitted TR-0516-49422-P, “Loss-of-Coolant Accident Evaluation Model,” Revision 0, (Agencywide Documents Access and Management System (ADAMS) Accession No. ML17004A202) to the U.S. Nuclear Regulatory Commission (NRC) for review and approval.

By letters dated November 27, 2019 and May 27, 2020, NuScale submitted, TR-0516-49422, Revisions 1 and 2, (ADAMS Accession Nos. ML19331B585 and ML20148T471, respectively). On June 19, 2020, NuScale supplemented TR-0516-49422, Revision 2, (ADAMS Accession Nos. ML20175A345). The NRC staff has evaluated TR-0516-49422, Revision 2, and found that it is acceptable for referencing licensing applications for the NuScale small modular reactor design to the extent specified and under the conditions and limitations delineated in the enclosed safety evaluation report (SER).

The NRC staff requests that NuScale publish the applicable version(s) of the SER listed above within 30 days of receipt of this letter. The accepted version of the TR shall incorporate this letter and the enclosed SER and add “-A” (designated accepted) following the report identification number.

CONTACT: Bruce M. Bavol, NRR/DNRL 301-415-6715 Z. Rad - 2 -

If the NRC staff’s criteria or regulations change, and its acceptability conclusion in the SER is invalidated, NuScale and/or the applicant referencing the SER will be expected to revise and resubmit its respective documentation; or submit justification for continued applicability of the SER without revision of the respective documentation.

After receiving the package with the “-A” version, the SER will be made available for public inspection through the publicly available records component of NRC’s ADAMS.

If you have any questions or comments concerning this matter, please contact Bruce Bavol at 301-415-6715 or via e-mail address at [email protected].

Sincerely,

/RA/

Anna H. Bradford, Director Division of New and Renewed Licenses Office of Nuclear Reactor Regulation

Docket No. 52-048

Enclosures: 1. TR-0516-49422 SER (Public) 2. TR-0516-49422 SER (Proprietary) cc: DC NuScale Power, LLC Listserv (w/o Enclosure 2)

Z. Rad - 3 -

SUBJECT: FINAL SAFETY EVALUATION FOR NUSCALE POWER, LLC TOPICAL REPORT TR-0516-49422, REVISION 2, “LOSS-OF-COOLANT ANALYSIS MODEL”

DISTRIBUTION: PUBLIC BBavol, NRR MDudek, NRR RPatton, NRR RidsNrrDnrl RidsNrrDss RidsEdoMailCenter RidsAcrsMailCenter RidsOgcMailCenter RidsNrrMailCenter RidsNrrLACSmithResource DC NuScale Power, LLC Listserv

ADAMS Accession Nos.: Pkg: ML20181A273 Letter: ML20181A270 Enclosure No. 1: ML20181A268 PUBLIC Enclosure No. 2: ML20181A269 PROP *via email NRR-106 OFFICE DNRL/NRLB: PM DNRL/NRLB: LA DNRL/NRLB: BC DNRL: D NAME BBavol CSmith* MDudek* ABradford* DATE 06/22/2020 06/22/2020 06/23/2020 06/22/2020 OFFICIAL RECORD COPY

U.S. NUCLEAR REGULATORY COMMISSION

SAFETY EVALUATION REPORT FOR TOPICAL REPORT TR-0516-49422, REVISION 2, May 2020

“LOSS-OF-COOLANT ACCIDENT EVALUATION MODEL”

Enclosure 1 TABLE OF CONTENTS

1.0 INTRODUCTION AND BACKGROUND ...... 1 2.0 REGULATORY BASIS FOR LOCA EM TOPICAL REVIEW ...... 2 2.1 Regulatory Requirements ...... 2 2.2 Regulatory Guide 1.203 ...... 5 2.3 NUREG-0800 Standard Review Plan ...... 7 3.0 NUSCALE LOCA EVALUATION METHODOLOGY SUMMARY...... 7 4.0 TECHNICAL EVALUATION ...... 9 4.1 Introduction and Scope ...... 9 4.2 Background...... 9 4.3 NuScale Power Module Description and Operations ...... 10 4.4 Phenomena Identification and Ranking ...... 11 4.5 Evaluation Model Description ...... 14 4.6 NRELAP5 Computer Code ...... 26 4.7 NRELAP5 Assessments ...... 36 4.8 Assessment of Evaluation Model Adequacy...... 48 4.9 Loss-of-Coolant Accident Calculations ...... 66 5.0 EVALUATION MODEL FOR INADVERTANT OPENING OF RPV VALVES ...... 74 5.1 Event Description and Classification ...... 74 5.2 Evaluation Model ...... 75 5.3 Accident Scenario Identification Process ...... 77 5.4 CHF Evaluation ...... 78 5.5 Experiential Data ...... 78 6.0 LIMITATIONS AND CONDITIONS ...... 90 7.0 CONCLUSION ...... 92

ii 1.0 INTRODUCTION AND BACKGROUND

Under the NRC Topical Report (TR) Program and by letter dated December 30, 2016, the applicant, NuScale Power, LLC (NuScale) submitted TR 0516 49422-P, “Loss-of-Coolant Accident Evaluation Model,” Revision 0, (Agencywide Documents Access and Management System (ADAMS) Accession No. ML17004A202), to the U.S. Nuclear Regulatory Commission (NRC) staff for review. The applicant supplemented its submittal by letter dated March 7, 2017 (ML17066A463). By letter dated April 27, 2017 (ML17116A063), the NRC informed NuScale of its acceptance of TR 0516 49422-P, Revision 0, for a detailed technical review. On November 27, 2019, (ML19331B585), NuScale submitted Revision 1 to TR 0516 49422-P and on May 27, 2020, (ML20148T471), NuScale Submitted Revision 2 to TR-0516-49422-P (hereafter referred to as the loss-of-coolant accident [LOCA] TR). By letter dated June 19, 2020, NuScale supplemented TR-0516-49422, Revision 2, (ML20171A731).

TR 0516 49422-P, “Loss-of-Coolant Accident Evaluation Model,” Revision 2, presents the NuScale’s evaluation model (EM) used to evaluate emergency core cooling systems’ (ECCS) performance in the NuScale Nuclear Power Module (NPM) for design basis LOCAs.

Additionally, Appendix B of TR-0516-49422, provides a description of a modified version of the LOCA EM that is used to evaluate the Inadvertent Opening of a Reactor Pressure Vessel (RPV) Valve (IORV) event. Section 5 includes details regarding the thermal limits evaluation during low flow, stagnation, and reverse flow, which occur during LOCAs and LOCA-like events. Specifically, it provides the bases to support: (1) the applicability of the [[ ]] critical heat flux (CHF) correlations (hereafter referred to as the high-flow and low-flow CHF correlations, respectively) for the analysis of the NPM, (2) the minimum CHF ratio (MCHFR) limit for each event, and (3) the range of applicability of these correlations. The TR also discusses the interfaces to the other analyses that assess the acceptance criteria not evaluated by the LOCA EM.

After the LOCA EM was developed, there was a design change to the NPM to ensure acceptable boron distribution during passive ECCS and decay heat removal system (DHRS) cooling modes. This included the addition of holes to the RPV riser and the addition of low RCS pressure emergency core cooling system actuation. Additionally, the TR was updated to describe the ECCS valves opening on low differential pressure between the RPV and CNV and the removal of the RPV level ECCS actuation. These changes were addressed in Revision 2 to TR-0516-49422-P.

NuScale stated that the LOCA EM was developed following the guidelines in the EM development and assessment process (EMDAP) of “Transient and Accident Analysis Methods,” Regulatory Guide (RG) 1.203 and that this model adheres to the applicable requirements under Title 10 of the Code of Federal Regulations (10 CFR) Part 50, Appendix K, “ECCS Evaluation Models,” and 10 CFR Section 50.46, “Acceptance Criteria for Emergency Core Cooling Systems for Light-Water Nuclear Power Reactors.” NuScale stated that multiple layers of conservatism are incorporated in the NuScale LOCA EM to ensure that a conservative analysis result is obtained. These conservatisms stem from an application of the modeling requirements of 10 CFR Part 50, Appendix K, and through a series of conservative modeling assumptions and modeling inputs.

This Safety Evaluation Report (SER) documents the results of the NRC staff’s in-depth technical evaluation of TR-0516-49422-P, “Loss-of-Coolant Accident Evaluation Model,” Revision 2, and 1

the NuScale EM used to evaluate the ECCS performance in the NuScale NPM. The NRC staff performed a review to determine the technical applicability of the thermal hydraulic methods and modeling techniques as described in TR-0516-49422-P, for evaluating ECCS core cooling performance for LOCA and LOCA-like events.

The applicant developed the NuScale LOCA Evaluation Methodology to evaluate ECCS performance for the NuScale NPM. The NuScale design is a small modular reactor designed to be deployed with up to 12 NPMs at a specific site. Each NPM is a light-water, integral pressurized water reactor (PWR) that is enclosed by a high-pressure containment vessel (CNV) immersed in a reactor pool coupled with passive safety-related ECCS. The NPM is designed to shut down and cooldown in the event of a LOCA. Each NPM has an independent nuclear steam supply system (NSSS) that includes a nuclear core, helical-coil steam generator (SG), integral pressurizer, strategically placed ECCS valves, and a compact, high-pressure steel CNV that contains the NSSS. Each NPM has a secondary system that includes a traditional steam- power conversion system including a steam turbine generator, condenser, and feedwater system. The integral small PWR design does not have large reactor system piping found in conventional PWRs, therefore the number and size of pipe ruptures that would result in a LOCA are significantly reduced. The NuScale LOCA EM evaluates potential breaks in the reactor coolant system (RCS) injection line, RCS discharge line, pressurizer spray supply line, and pressurizer high point vent line. The RCS injection line is supplied by the chemical and volume control system (CVCS) and the discharge line returns to the CVCS. In addition, the applicant extended the EM to evaluate the design basis events resulting from an IORV. During normal operation, flow through the reactor is driven by natural circulation resulting from the thermal driving head produced by the temperature difference between the core and the heat sink afforded by the SGs. Natural circulation flow increases reliability that ECCS will successfully initiate recirculation flow by eliminating primary coolant pumps that can fail or lock up. NuScale designed the NPM so that there is no core uncovery or heatup for a design-basis LOCA.

2.0 REGULATORY BASIS FOR LOCA EM TOPICAL REVIEW

The NRC staff has reviewed the LOCA EM described in TR-0516-49422-P, Revision 2, entitled “Loss-of-Coolant Accident Evaluation Model,” to determine whether this methodology is acceptable for performing LOCA analyses and IORV analyses. The NRC staff also reviewed the basis for applying the NRELAP5 code to predict certain highly-ranked phenomena that govern peak containment pressure analyses. The methodology for calculating the peak containment pressure and temperature performance is contained in the Containment Response Analysis Methodology (CRAM), TR-0516-49084, Revision 2 (ML19330F387), which is an extension of the LOCA TR methodology. This section of the SER describes the regulatory basis and supporting regulatory and guidance documents that the NRC staff uses to determine whether the methodology described in LOCA EM TR-0516-49422-P, Revision 2, is acceptable for LOCA and IORV analyses of the NuScale NPM design.

2.1 Regulatory Requirements

The requirements under 10 CFR 50.46 and 10 CFR Part 50, Appendix K, present the acceptance criteria for ECCS for light water nuclear power reactors and the required and acceptable features of the EMs employed. The NuScale LOCA EM is based on meeting the conservative 10 CFR part 50, Appendix K rule per 10 CFR 50.46(a)(1)(ii).

2.1.1 10 CFR 50.46 ECCS and Appendix K to 10 CFR Part 50 Requirements

The ECCS Rule at 10 CFR 50.46, “Acceptance Criteria for Emergency Core Cooling Systems for Light-Water Nuclear Power Reactors,” requires in 10 CFR 50.46(a)(1)(i) that each PWR fueled with uranium oxide pellets within cylindrical zircaloy or ZIRLO cladding must be equipped with an ECCS and that ECCS performance must be evaluated for the most severe postulated accident.

ECCS Analysis Method

The regulations at 10 CFR 50.46(a)(1)(i) requires that “ECCS cooling performance must be calculated in accordance with an acceptable evaluation model,” and provides for two options, as mentioned above, for acceptable EM analytical techniques and methods: realistic or conservative.

Accordingly, 10 CFR 50.46(a)(1)(ii) describes an EM of the second category as a method that conservatively describes the behavior of the reactor system during a loss-of-coolant accident. and that such an EM “may be developed in conformance with the required and acceptable features” of Appendix K, “ECCS Evaluation Models,” to 10 CFR Part 50.

Furthermore, 10 CFR 50.46(c)(2) defines an EM as the calculational framework for evaluating the behavior of the reactor system during a postulated LOCA. An EM includes one or more computer programs and all other information necessary for applying the calculational framework to a specific LOCA (the mathematical models used, the assumptions included in the programs, the procedure for treating the program input and output information, the parts of the analysis not included in the computer programs, values of parameters, and all other information necessary to specify the calculational procedure).

10 CFR Part 50 Appendix K

The regulation at 10 CFR 50.46(a)(1)(ii) provides that an EM may be developed in conformance with the required and acceptable features that include specific conditions to be met as defined by Appendix K to 10 CFR Part 50.

Since the NuScale EM does not present any LOCA predictions that produce core uncovery and exposure of the active fuel region to steam cooling, neither fuel cladding damage or cladding oxidation is calculated, eliminating the potential of several 50.46 criteria from being exceeded. As such, the NuScale LOCA TR includes the provision that, “A feature “excluded” from the EM means that 10 CFR 50, Appendix K, directly requires the feature, without condition on the presence of a process or phenomena, but that the feature is not relevant to the NuScale LOCA EM. Table 2-2 technically justifies the exclusion of such feature from the model. However, an applicant or licensee referencing this report will be required to address regulatory compliance with 10 CFR 50.46 and 10 CFR 50, Appendix K (e.g., by seeking an exemption from that required feature).”

As stated in the applicability section of the TR, the review of the NuScale EM does not apply to conditions where the liquid level recedes below the top elevation of the core active fuel region. For this reason, the NuScale EM does not contain post-CHF clad damage models and metal water reaction methodologies that would normally accompany a PWR LOCA TR submitted for the NRC staff’s review. However, should the predicted liquid level calculated using the NuScale 3

EM ever recede below the top elevation of the core, those conditions are beyond the capability of this methodology and modification to this EM would need to be submitted for the NRC staff’s review and approval.

These modifications would include, but are not be limited to, post-CHF heat transfer models, fuel pin models that incorporates clad swelling, rupture and, oxidation, and calculation of the metal-water reaction rate using the Baker-Just Correlation.

Part II, “Required Documentation,” of Appendix K, “ECCS Evaluation Models,” to 10 CFR Part 50, sets forth the EM documentation requirements for the required analyses as well as the need for additional sensitivity studies and comparisons of the EM to experimental data.

The regulation at 10 CFR 50.46(a)(1)(i), requires, in part, that “ECCS cooling performance must be calculated in accordance with an acceptable evaluation model and must be calculated for a number of postulated loss-of-coolant accidents of different sizes, locations, and other properties sufficient to provide assurance that the most severe postulated loss-of-coolant accidents are calculated.”

ECCS Performance Criteria

The regulation at 10 CFR 50.46(a)(1)(i), requires, in part, that the ECCS calculated cooling performance following postulated LOCAs conforms to the criteria set forth in 10 CFR 50.46(b). This regulation defines the criteria for the calculated ECCS cooling performance during postulated LOCAs in 10 CFR 50.46(b)(1) through 10 CFR 50.46(b)(5) as follows:

(1) Peak Cladding Temperature. The calculated maximum fuel element cladding temperature shall not exceed 2,200 degrees Fahrenheit (ºF) (1,477.59 K or 1,204.44 degrees Celsius (ºC)).

(2) Maximum Cladding Oxidation. The calculated total oxidation of the cladding shall nowhere exceed 0.17 times the total cladding thickness before oxidation. This is based on the Baker-Just equation.

(3) Maximum Hydrogen Generation. The calculated total amount of hydrogen generated from the chemical reaction of the cladding with water or steam shall not exceed 0.01 times the hypothetical amount that would be generated if all of the metal in the cladding cylinders surrounding the fuel, excluding the cladding surrounding the plenum volume, were to react.

(4) Coolable Geometry. Calculated changes in core geometry shall be such that the core remains amenable to cooling.

(5) Long-Term Cooling. After any calculated successful initial operation of the ECCS, the calculated core temperature shall be maintained at an acceptably low value and decay heat shall be

4 removed for the extended period of time required by the long-lived radioactivity remaining in the core.

2.1.2 10 CFR 50, Appendix A, “General Design Criteria for Nuclear Power Plants”

In TR-0516-49422, the applicant requests approval for two CHF models to be used in accordance with the analysis methodologies described in the TR. The CHF correlations, and their respective limits, are used to evaluate whether fuel cladding integrity is maintained during LOCA and LOCA-like events. Thus, approved CHF correlations and associated methodologies are used to establish a partial basis for compliance with the following general design criteria (GDC) in 10 CFR Part 50, Appendix A, “General Design Criteria for Nuclear Power Plants.”

• GDC 10, “Reactor Design,” which requires that the reactor core and associated coolant, control, and protection systems be designed with appropriate margin to assure that specified acceptance fuel design limits are not exceeded during any condition of normal operation, including the effects of anticipated operational occurrences (AOOs).

• GDC 19, “Control Room,” and 10 CFR 52.47(a)(2)(iv) as they relate to the evaluation and analysis of the radiological consequences from postulated accidents.

• GDC 35, “Emergency Core Cooling,” as it relates to demonstrating that the ECCS would provide abundant emergency core cooling to satisfy the ECCS safety function of transferring heat from the reactor core following any loss of reactor coolant at a rate that: (1) fuel and clad damage that could interfere with continued effective core cooling would be prevented, and (2) clad metal-water reaction would be limited to negligible amounts.

2.2 Regulatory Guide 1.203

RG 1.203, “Transient and Accident Analysis Methods,” provides guidance for developing and evaluating EMs for accident and transient analyses. Section D, “Implementation,” states that the guide is approved for use as an acceptable means of complying with the NRC regulations and for evaluating submittals of “new or modified EMs proposed by vendors or operating reactor licensees that, in accordance with 10 CFR 50.59, require NRC staffs review and approval.”

The LOCA EM is a deterministic analysis approach NuScale developed considering the requirements of 10 CFR 50.46 and 10 CFR Part 50, Appendix K. The LOCA TR states that the approach to the development of the model follows RG 1.203, “Transient and Accident Analysis Methods.” Within RG 1.203, the Phenomena Identification and Ranking Table (PIRT), is identified as a key requirement for EM development. Section 4 of the NuScale LOCA EM TR documents the PIRT that NuScale developed for the NPM. Section 4.4 of this SER provides the NRC staff’s review of this PIRT.

2.2.1 Evaluation Model Concept

In accordance with 10 CFR 50.46(c)(2), RG 1.203 states that the EM constitutes the calculational framework for evaluating the behavior of the reactor system during a postulated transient or a design-basis accident. As such, the EM may include one or more computer programs, special models, and all other information needed to apply the calculational framework 5 to a specific event, such as procedures for treating the input and output information, specification of those portions of the analysis not included in the computer programs for which alternative approaches are used, or all other information needed to specify the calculational procedure. It is the entirety of an EM that ultimately determines whether the results comply with applicable regulations and therefore the development, assessment, and review processes must consider the entire EM. Most EMs used to analyze the events in SRP Chapter 15, “Transient and Accident Analysis,” rely on a systems code that describes the transport of fluid mass, momentum, and energy throughout the RCSs. The LOCA EM uses the NuScale NRELAP5 systems analysis computer code, which is developed from the Idaho National Laboratory (INL) RELAP5-3D computer code.

2.2.2 Evaluation Model Development and Assessment Principles

RG 1.203 defines the following six basic principles as important to follow in the EMDAP:

(1) Determine requirements for the EM. (2) Develop an assessment base consistent with the determined requirements. (3) Develop the EM. (4) Assess the adequacy of the EM. (5) Follow an appropriate quality assurance (QA) protocol during the EMDAP. (6) Provide comprehensive, accurate, up-to-date documentation. RG 1.203 discusses the NRC staff’s regulatory position, which provides guidance concerning methods for calculating transient and accident behavior. Part C of RG 1.203, provides guidance on aspects of an EMDAP that are related to the basic principles identified above and offers additional guidance.

Regulatory Position 1, EM Development and Assessment Process (EMDAP)

RG 1.203 identifies four basic elements developed to describe an EMDAP. The elements correspond to the first four EMDAP basic principles and provide guidance in twenty individual steps. In addition, Regulatory Position 1 includes requirements for reaching an adequacy decision. The basic elements of Regulatory Position 1 are identified below.

Element 1: Establish Requirements for EM Capability Element 2: Develop Assessment Base Element 3: Develop EM Element 4: Assess EM Adequacy Decision Regulatory Position 2, Quality Assurance

RG 1.203 discusses QA during development, assessment, and application of an EM and the requirements of Appendix B, “Quality Assurance Criteria for Nuclear Power Plants and Fuel Reprocessing Plants,” to 10 CFR Part 50.

Regulatory Position 3, Documentation

RG 1.203 provides guidance on the requirements to document the development of LOCA EMs.

6 Regulatory Position 4, General Purpose Computer Programs

RG 1.203 provides guidance on development of general-purpose transient analysis computer programs designed to analyze a number of different events for a wide variety of plants. Specifically, Regulatory Position 4 states that “application of the EMDAP should be considered as a prerequisite before submitting a general-purpose transient analysis computer program for review as the basis for EMs that may be used for a variety of plant and accident types.”

Regulatory Position 5, Graded Approach to Applying the EMDAP Process

RG 1.203 provides guidance on the extent to which the full EMDAP should be applied for a specific application based on the following four EM attributes:

(1) Novelty of the revised EM compared to currently accepted models. (2) Complexity of the event being analyzed. (3) Degree of conservatism in the EM. (4) Extent of any plant design or operational changes that would require reanalysis. Appendix A of RG 1.203, “Additional Considerations in the Use of this RG for ECCS Analysis,” describes uncertainty determination and provides guidance for best-estimate LOCA analyses. Appendix A of RG 1.203 refers to NUREG-0800, “Standard Review Plan for the Review of Safety Analysis Reports for Nuclear Power Plants: LWR Edition,” (SRP), Sections 15.6.5, “Loss-of-Coolant Accidents Resulting From Spectrum of Postulated Piping Breaks Within the Reactor Coolant Pressure Boundary,” and 15.0.2, “Review of Transient and Accident Analysis Method.”

2.3 NUREG-0800 Standard Review Plan

NUREG-0800, “Standard Review Plan for the Review of Safety Analysis Reports for Nuclear Power Plants: LWR Edition,” (SRP), Section 15.0.2, “Review of Transient and Accident Analysis Methods,” is the companion SRP section for RG 1.203.

SRP Section 15.6.5, “Loss-of-Coolant Accidents Resulting from Spectrum of Postulated Piping Breaks within the Reactor Coolant Pressure Boundary,” Revision 3, describes the review scope, acceptance criteria, review procedures, and findings relevant to ECCS analyses.

3.0 NUSCALE LOCA EVALUATION METHODOLOGY SUMMARY

The NuScale Power Module has several unique features that required the NRC staff to perform detailed reviews of the NuScale LOCA EM to determine whether this methodology is adequate. The NuScale design is a small modular PWR that relies on natural circulation during normal plant operation and uses a unique high-pressure containment as an integral part of the ECCS to keep the reactor core covered with the collapsed liquid level (CLL) above the top of the active core through all potential LOCA events, as shown in NuScale LOCA TR Figures 3-1, “A singular NuScale Power Module during normal operation,” and 3-2, “Schematic of NuScale Power Module decay heat removal system and emergency core cooling system during operation.”

During a NuScale NPM LOCA, the high-pressure water and steam leaving the RPV is contained in the CNV. The CNV has a design pressure of 1050 pounds per square inch absolute (psia)

7 and is designed to enable the ECCS system to return cooled RCS liquid to the downcomer to prevent core uncovery during design basis LOCAs. During a LOCA, the five ECCS valves, three Reactor Vent valves (RVVs) and two Reactor Recirculation Valves (RRVs), receive a signal to open. However, the valves are blocked from opening by the Inadvertent Actuation Block (IAB) Valve until the pressure differential between the RPV and CNV drops below the IAB threshold. Once these valves open, the RPV and CNV pressures equalize within about 30 seconds. After this pressure equalization, steam generated inside the RPV from decay heat and stored energy, exits the RPV through the RVV, condenses on the inside of the CNV wall, and is returned from the CNV to the RPV through the RRVs.

Because of the unique features of the NuScale NPM containment (CNV) design and the NuScale ECCS system, the NRC staff’s review of the NuScale LOCA EM TR focused particular attention on the ability of the NuScale LOCA EM to assess the following design issues and phenomena:

• The capability to predict the CLL in the RPV so that the NuScale power module maintains the CLL in the RPV above the reactor core during the design basis event of a LOCA. • The capability to predict Critical Heat Flux Ratio (CHFR) so that the NuScale power module maintains the CHF margin during the design basis event of a LOCA. • The applicability of NRELAP5 computer code to perform peak containment pressure so that the CNV of NuScale power module absorbs heat energy at a rate sufficient to maintain CNV pressure within design limits and to transfer heat energy from the RPV to the water pool outside the CNV during the design basis event of a LOCA • Ensuring that the LOCA pipe break spectrum methodology included all susceptible RPV penetrations. In addition, the NRC staff’s review of the EM TR focused particular attention on the capability of the NuScale NRELAP5 computer code to accurately model the tests performed at the NPM scaled model NuScale Integral System Test Facility (NIST-1) facility and to confirm that the geometric dimensions and operating conditions of NIST-1, adequately represent the NPM full plant.

Because the NuScale design relies on maintaining a CLL above the top of the reactor core, the NRC staff’s evaluation of the NuScale LOCA Evaluation Methodology is limited to the consideration of the conservative assumptions and modelling assumptions to determine that this design objective is adequately modeled. The determination to support the Design Certification that the CLL remains above the top of the core, is documented as part of the review of the NuScale Design Certification Application.

The NRELAP5 computer code, Version 1.4 (ML17066A463 and ML19162A086), was submitted as the systems analysis computer code for the NuScale LOCA Evaluation Methodology. NuScale’s primary changes to the INL RELAP5-3D version included implementation of a new helical-coil SG (HCSG) component and the addition of new containment condensation models to describe the unique design features of the ECCS operation of the NPM. In addition to the use of NRELAP5 for evaluating LOCAs, Appendix B of the LOCA TR extends the EM and methodology to analyze events described in SRP Section 15.6.1, “Inadvertent Opening of a PWR Pressurizer Pressure Relief Valve,” and SRP Section 15.6.6, “Inadvertent Operation of the Emergency Core Cooling System (ECCS) event.” These inadvertent RPV 8

valve events are classified as AOOs so the acceptance criteria are also slightly different than the Section 15.6.5 LOCA acceptance criteria. The event progression is essentially that of a LOCA which results in blowdown of the RCS inventory into the CNV and can be a steam region release or liquid space discharge.

4.0 TECHNICAL EVALUATION

This section of the SER summarizes and evaluates the information in each section of the TR against the regulatory requirements for that section. The Limitations and Conditions on the review of TR-0516-49422-P, are discussed in detail below and summarized in Section 6. The conclusions from the review are discussed in detail below and summarized in Section 7 of this SER.

In addition, the NRC staff conducted audits of information provided by the applicant in support of the NRC staff’s review of the TR. These audits are referenced throughout this SER. Unless otherwise noted, details of these audits are available in audit reports at ADAMS Accession Nos. ML20010D112, ML20034D464, and ML20160A250, which provide summaries of the audits and the information examined during them.

4.1 Introduction and Scope

Section 1.1 of TR-0516-49422-P, states that the purpose of the NuScale EM is to evaluate ECCS performance in the NPM for design-basis LOCAs and requests approval to use the methodology to perform such analyses. NuScale stated that its LOCA EM follows the guidance provided in “Transient and Accident Analysis Methods,” RG 1.203 and satisfies the applicable requirements of "ECCS Evaluation Models," 10 CFR Part 50, Appendix K. TR-0516-49422-P provides a description of the methodology used by NuScale for LOCA analyses and this methodology is reviewed in this SER for compliance with applicable regulatory criteria. However, TR-0516-49422-P does not provide any final licensing analyses and this review of TR-0516-49422-P does not evaluate the acceptability of the NuScale NPM or provide any conclusions on the acceptability of the NuScale design.

Further, the LOCA TR provides support for other analyses including:

1. events as described in TR-0516-49416-P, “Non-Loss of Coolant Accident Methodology”;

2. containment peak pressure analysis as described in Technical Report TR-0516-49084-P, “Containment Response Analysis Methodology”;

3. long term cooling as described in Technical Report, TR-0919-51299-P, “Long-Term Cooling Methodology”; and

4. IORV valves, including ECCS valves as described in Appendix B of the LOCA TR, “Evaluation Model for Inadvertent Opening of RPV Valves.” 4.2 Background

Section 2 of the NuScale TR provides a description of how the NuScale LOCA EM conforms with the EMDAP guidance in RG 1.203. Additionally, NuScale stated that other provisions of

RG 1.203 related to establishing an appropriate QA program (QAP) and providing comprehensive, accurate, up-to-date documentation are described outside of the LOCA EM TR. QA requirements are included in “NuScale Topical Report: Quality Assurance Program Description for the NuScale Power Plant,” NP-TR-1010-859-NP-A. The NRC staff reviewed the QA requirements and documented its approval in its SER (ML16347A405). Further, the NRC staff inspected NuScale’s design control process and code development procedures. These inspections are documented in inspection reports dated October 7, 2017 (ML15268A186) and July 24, 2017 (ML17201J382).

NuScale indicated that the NPM is designed to reduce the consequences of design-basis LOCAs compared to existing PWRs for which 10 CFR Part 50, Appendix K was developed. Consequently, many of the phenomena that are the subject of 10 CFR Part 50, Appendix K requirements are not encountered in design-basis NPM event LOCAs, meaning that these phenomena have been eliminated by design, therefore, a number of the Appendix K requirements are satisfied by design rather than by analysis. NuScale has therefore limited the methodology to only pre-CHF heat transfer regimes and specified that an applicant or licensee seeking to reference the LOCA TR, must demonstrate regulatory compliance for these Appendix K requirements, which could include seeking an exemption. The requirements are included in Table 2-2, “10 CFR 50 Appendix K required and acceptable features compliance,” of the TR. This is reflected in Section 6.0, Limitations and Conditions, of this report. The staff’s review in this SER is therefore limited to pre-CHF heat transfer regimes.

The NuScale LOCA EM TR provides a summary of the NRELAP5 code modifications and modeling features added by NuScale to address the unique features and phenomena of the NPM design and states that the EM is consistent with the applicable requirements of 10 CFR Part 50, Appendix K and the Three Mile Island (TMI) Action Items applicable to the NuScale NPM as described in the Design-Specific Review Standard for NuScale, Section 15.6.5. The NRC staff’s review of the CHF correlations used in NRELAP5 are contained in Section 4.6.

4.3 NuScale Power Module Description and Operations

Section 3 of the NuScale LOCA EM TR provides a brief description of the NPM and a brief summary of NPM operation.

4.3.1 General Plant Design

Features of the NuScale plant design that are unique compared with existing operating PWR plants include:

* Reduced reactor core size.

* Natural circulation reactor coolant flow (i.e., no reactor coolant pumps).

* An integrated HCSG and a pressurizer inside the RPV that eliminates piping to connect the SG or pressurizer with the reactor.

* A safety-related ECCS system that does not require electrical power and does not use ECCS pumps.

* Primary fluid in the SGs flow on the outside of the tube surface, and two-phase flow of the secondary flow contained inside of the tubes.

* A high-pressure steel CNV immersed in a water-filled pool that is integral to the ECCS capability to provide for emergency cooling.

In TR Section 3.1, “General Plant Design,” NuScale describes how the NPM uses natural circulation to provide reactor core coolant flow without electrically power reactor coolant pumps. This section of the LOCA TR also describes the design of the NPM HCSG. As discussed in the LOCA TR, each NPM has a dedicated ECCS, CVCS, and DHRS. The NRC staff reviewed the summary of its general plant design in TR-0516-49422-P and found that it provided sufficient description of the design to support the methodology description.

4.3.2 Plant Operation

In TR Section 3.2, NuScale provides a brief description of NPM operation including systems modeled in the LOCA evaluation methodology. The NRC staff reviewed the plant operation summary in TR-0516-49422-P and found that it provided a sufficient description to support the methodology description.

4.3.3 Safety-Related System Operation

In Section 3.3 of the TR, NuScale describes operation of the safety-related systems and components, including the ECCS, the NuScale Module Protection System, DHRS and the containment isolation valves. The NRC staff reviewed the safety-related system operation summary in TR-0516-49422-P and found that it provided a sufficient description to support the methodology description.

The ECCS is a two-phase natural circulation system that is designed to maintain a water supply to the core during its operation in a LOCA scenario. The RPV and CNV geometry is designed such that ECCS actuation results in a CLL in the RPV that is generally significantly above the top of the core. The ECCS is actuated on high CNV level or low RCS pressure interlocked with RCS hot temperature and CNV pressure or at 24 hours with loss of AC power. If the ECCS is not already open by previous mechanisms, there is also a low differential RPV to CNV pressure feature due to the valve spring that can open the valves at about 15 psid.

The DHRS is a passive safety-related system that uses boiling condensation loop flow to remove heat from the RCS through the SG and reject heat to the reactor pool through the DHRS condensers. The DHRS is composed of two DHRS trains associated with the NPM SG and each train is designed with the capability to independently remove 100 percent of decay heat.

4.4 Phenomena Identification and Ranking

As discussed in Section 4 of the NuScale LOCA EM TR, NuScale developed the NPM PIRT in stages. NuScale developed its original PIRT in 2008 and updated this PIRT in 2013 and 2015. NuScale used the 2015 final PIRT as the basis for the presentation given in Chapter 4 of the LOCA TR. However, NuScale only documented phenomena and processes of high importance in Section 4 of the NuScale LOCA EM TR. Therefore, the NRC staff reviewed both the LOCA

TR and the NuScale 2015 PIRT. The 2015 NuScale PIRT provided the rankings for all four PIRT importance categories (high, medium, low, and inactive).

NuScale’s first step in the PIRT development was to select the panel to support the PIRT review and to examine the qualifications of the PIRT board members to assure that they were qualified. NuScale’s second step of PIRT development was to obtain an agreement between the NuScale staff and the PIRT panel on the accident scenarios and figures of merit (FOMs) identified by NuScale and the PIRT panels. NuScale’s third step was to review of each of the phenomena and processes identified and ranked in the PIRT to determine the approximate fidelity of the rating assigned. In this process, NuScale did not require unanimous agreement on the reasons given for the PIRT ranking. However, NuScale did use a process to assure that no phenomena or process of high importance was missing and that the rankings of medium and low importance were reasonable.

The NRC staff reviewed the PIRT panel membership and the qualifications of the NuScale 2008, 2013 and 2015 PIRT panels as provided in Section 4 of the LOCA EM TR and the 2015 PIRT report that the staff audited (ML20010D112). NuScale provided the list of PIRT panel members in the LOCA TR. However, NuScale only provided panel member qualifications for the 2015 PIRT panel in the 2015 report. Based both on the NRC staff’s knowledge of the panel members listed and the NRC staff’s examination of the qualifications provided in the 2015 report, the NRC staff confirmed that all panel members are highly regarded members of the nuclear community with extensive experience in the industry, a research institution, or nuclear academia.

4.4.1 NuScale Loss-of-Coolant Accident Scenarios

NuScale discussed LOCA accident scenarios in LOCA EM TR Section 4.2, which it states are consistent with 10 CFR 50.46(c)(1). NuScale stated that it considered breaks of various sizes, types, and orientations in piping connected to the RPV. Because the NuScale NPM design eliminates most primary coolant piping, breaks are limited to the RCS injection and discharge lines, the pressurizer spray supply line, and the pressurizer high point vent line. [[

]] The TR provides a description of the progression of each LOCA scenario and divides the scenarios into two phases: LOCA blowdown (1a) and ECCS actuation to the time when stable long-term recirculation flow is established (1b).

The NRC staff evaluated the LOCA scenarios selected for the PIRT discussions and determined that the only LOCAs that can occur are those for penetrations of the RPV that pass into or through the CNV and originate within the RCS. These penetrations are few and small in cross- sectional area. Therefore, the NRC staff agreed that the large break LOCA scenarios for conventional pressurized reactor designs that circulate reactor coolant through large pipes outside of the RPV, are not applicable for the NuScale NPM.

The NRC staff found that NuScale’s PIRT phenomenon selection of steam breaks from [[ ]] The NRC staff notes that for the NuScale design, breaks from the three RVV nozzles and two RRV flanges are excluded in this TR as break locations and are the subject of a separate evaluation in the DCD, Chapter 3, “Design of Structures, Systems, Components and Equipment,” regarding their exclusion from consideration as break locations. Additionally,

NuScale has evaluated inadvertent openings of an RRV or RVV as AOOs. The NRC staff’s review of the NRELAP5 code for evaluation of an inadvertent opening of an RRV or RVV is discussed in Section 5 of this SER.

The NRC staff determined that the NuScale LOCA accident scenarios selected as the basis for their PIRT process are acceptable for establishing the ranking phenomena that must be considered in the LOCA EM because they consider the applicable features of the design. 4.4.2 Figures of Merit

In TR Section 4.3, NuScale discusses the FOMs selected for its LOCA EM, which are primarily CHFR and CLL above the top of active fuel (TAF). According to NuScale, the NPM retains sufficient water in the RPV such that the core will not be uncovered during any LOCA scenario. Therefore, peak clad temperature is not a FOM for the NuScale PIRT process. Instead, the CHFR is an important FOM used to demonstrate that the fuel clad does not reach the point of CHF where significant heat up of the cladding could occur. NuScale also states that maintaining the CLL above the core is an additional LOCA analysis FOM as it demonstrates that there is an adequate supply of liquid water available to preclude CHF in the core. As a result of the scenarios selected, the FOMs for the PIRT considerations are the fuel rod CHF value and the requirement to maintain CLL above the top of the active core fuel.

The NRC staff finds for the NuScale LOCA EM FOMs that: (1) the core fuel rods do not experience CHF and that (2) the CLL in the RPV remains above the core at all times during all LOCA scenarios, shows conservatism, and are acceptable FOMs for the NuScale LOCA EM, particularly in light of the fact that the EM is limited to pre-CHF heat transfer because these FOMs are more restrictive than that required by 10 CFR 50.46(b).

Because the NuScale LOCA EM includes only pre-CHF phenomena, credible LOCA break scenarios must not produce core uncovery or thermal-hydraulic conditions which result in exceeding the CHF limits. The NRC staff has determined that the validity of the NuScale LOCA EM model is limited to LOCA analyses that do not reach core uncovery and where the heat flux for fuel cladding remains below the CHFR limit. This Limitation is reflected in Section 6.0, “Limitations and Conditions,” of this SER.

NuScale stated that to ensure ECCS performance, CNV must be maintained and intact during all postulated accident scenarios. Therefore, the CNV must be kept below CNV design pressure and temperature design limits to ensure compliance with 10 CFR 50.46 criteria. However, NuScale stated that the limiting peak containment pressure and temperature are calculated with a different methodology from that described in the LOCA EM TR. NuScale developed methodology for ensuring that containment integrity is presented in TR-0516-49084-P, “Containment Response Analysis Methodology Technical Report” (ML19330F387).

Since the NuScale LOCA EM depends on showing that no credible LOCA scenario or break can result in a loss of containment integrity, the NRC staff has determined that the validity of the NuScale LOCA EM is limited to LOCA analyses that result in CNV temperature and pressure that remain below respective design limits for all LOCA events as required for the ECCS system to maintain sufficient liquid water in the CNV to ensure that the CLL remains above the top of the reactor core. This is reflected in Section 6.0, “Limitations and Conditions,” of this SER.

4.4.3 PIRT Rankings

Sections 4.4 through 4.7, of the LOCA EM TR discuss the results of the NuScale LOCA EM PIRT process and provide a list of High-Ranked Phenomena and the Phenomena Identification and Ranking Summary Table.

The NuScale PIRT panel identified phenomena and processes that could occur during a NPM LOCA, ranked relative importance of each, and assessed the knowledge level on each. Relative importance was ranked as High, Medium, Low, or inactive (not present or negligible). Knowledge level was divided into well known, known, partially known, or very limited. Finally, the portion of the NPM for which the phenomena or process was ranked, was identified.

In LOCA EM Table 4-4 “High-Ranked phenomena,” NuScale provided the listing of the findings of its final PIRT for phenomena ranked of high importance as to position within the system where the phenomena are important, timing of importance during the accident, and knowledge level. NuScale assessed only the high ranked phenomena and processes in its LOCA EM TR. To assess medium to low ranked phenomena, the NRC staff reviewed the NuScale 2015 PIRT report. The NRC staff’s review did not identify any low to medium ranked phenomena or processes that should have been ranked high. Therefore, the NRC staff finds that the NuScale LOCA EM table for high ranking phenomena, is acceptable.

The NRC staff also reviewed TR Section 4.6.1 of the LOCA TR, “Discussion of Phenomena Ranked High Importance.” The NRC staff found that the rationale for the rankings in Table 4-4 is reasonable and appropriate but not always comprehensive. For example, the first entry in Section 4.6.1 [[

]] The NRC staff agrees that mass and energy release are a key factor but not the only key factor. The NRC staff found that the rate of mass and energy release is more important than the total amount of release in the determination of the CNV pressure and the flow changes in the RCS. However, several key important highly ranked phenomena included [[

]]. Staff subsequently agreed that assessment of rate is implied in the assessment of choke and unchoked flow. Therefore, the NRC staff finds that the conclusions and the PIRT phenomena selections and knowledge level rankings are appropriate as a basis for the NuScale LOCA EM. 4.5 Evaluation Model Description

The NRC staff reviewed the NuScale LOCA EM description provided in Section 5 of the NuScale LOCA EM TR to determine whether the analysis model described in Section 5 is suitable for performing LOCA safety analysis with NRELAP5. Section 5 describes the NPM nodalization and modeling input options selected by NuScale for each NPM component and provides NuScale’s rationale for each choice. Section 5 also provides NuScale’s justification for the boundary and initial conditions selected by NuScale, for the model. In addition, NuScale described the LOCA break spectrum selected by NuScale. NuScale stated that its NPM LOCA modeling used, is consistent with the Separate Effects Tests (SET) and Integral Effects Tests

(IET) assessments used by NuScale to validate the NRELAP5 code for its application to LOCA and Non-LOCA analysis.

4.5.1 NRELAP5 Loss-of-Coolant Accident Model for the NuScale Power Module

In LOCA EM TR Section 5.1, NuScale stated that its unique design features of the NPM allowed NuScale to use a simplified modeling approach to predict and evaluate consequences of postulated LOCAs.

The NRC staff’s review of the NuScale LOCA EM is based on these key NuScale design assumptions. In Section 5, NuScale stated that in the event of a LOCA, these unique NPM design features result in a simple, predictable transient progression, that can be explained by a standard mass and energy balance over the RPV and CNV considering:

• Choked and unchoked flow through the break and then ECCS flow via valves between RPV and CNV,

• Core decay heat generation and RCS stored energy release, and

• Heat transfer between the CNV and the reactor pool that is characterized by steam condensation at the CNV inside surface and free convection at the CNV outside surface to the reactor pool. The NRC staff reviewed the adequacy of the NRELAP5 modeling of these design features and determined that the modeling approach is adequate to evaluate the FOM. The NRC staff found that NPM modeling developed adequately represent the key components and the key phenomena expected to occur during a LOCA.

NuScale Power Module NRELAP5 Model

The NuScale LOCA EM covers key components of the NPM participating in a LOCA. These key components include the following:

1. RPV with internals:

a. Lower plenum b. Reactor core c. Riser including the riser upper plenum d. Upper and lower downcomer e. Pressurizer

2. Containment (CNV)

3. SG secondary side

4. Reactor pool

5. ECCS valves

6. Postulated break locations

7. RPV internal heat structures and heat structures between components (i.e., RPV to the CNV to the reactor pool).

8. Riser holes located at approximately the midpoint of the SG (note that these holes were not incorporated into the models used in producing the example results described in the TR, but is in the model the staff is approving for use in this evaluation through limitation and condition 5 in Section 6 of this report. This is addressed later in this section of the report)

The nodalization diagram of these key components is shown in Figure 5-1, “Noding diagram of NRELAP5 loss-of-coolant accident input model for NuScale Power Module,” of the LOCA EM TR.

In LOCA EM TR Section 5.1.1, “General Model Nodalization,” NuScale stated that the NRELAP5 RCS noding was developed to provide appropriate resolution of fluid volumes as a function of elevation to account for natural circulation flow during NPM operations and to calculate the draining of fluids into the lower RCS volumes and circulation between the containment, the downcomer, and the core when the ECCS system is activated later in the LOCA.

The NRC staff finds that the LOCA model adequately represents the important components and phenomena required for evaluating LOCA scenarios for the NPM.

Section 5.1.2, “Reactor Coolant System,” of the LOCA TR describes the modeling of each of the RCS components. The NRC staff’s findings relative to the modeling of those components is described below.

Downcomer, Lower Plenum and Riser

The NRC staff reviewed the description of the modeling of these three regions and finds that the model is acceptable because the model correctly preserved the volume, elevation changes in these three regions and incorporates conservative loss factors to maximize core bypass flow, the flow from the downcomer and into the reactor core.

After the TR was initially developed, a design change added holes to the RPV riser. This design change was not incorporated into the model used in the examples in this TR and is not included in the noding diagram in Figure 5-1. However, the staff determined that these holes would not significantly impact the example calculations used to demonstrate an acceptable EM. Additionally, limitation and condition 5 in Section 6 of this report requires the use of NPM model Revision 3, which includes the riser holes.

Reactor Core Model

The NuScale RELAP5 model uses three axial channels in the reactor core to calculate reactor coolant flow through the core, including hot, average and bypass channels.

During development of the LOCA EM, NuScale recognized that flow reversal may occur within the core bypass at low or stagnant flow conditions. Therefore, NuScale selected sufficient axial nodes to account for the hydrostatic head in these three parallel reactor core channels. NuScale modeled the unheated sections of the hot and average core channels with single nodes below and above the core heated length. The core model includes form losses for top 16 and bottom nozzles and grid spacers with appropriate hydrodynamic volumes based on fuel vendor data. NuScale modeled the core flow channels with individual NRELAP5 PIPE components. The crossflow between hot and average fuel channels was not credited to conservatively prevent hot and average channel fluid streams from mixing, which the NRC staff considers conservative.

The NPM model sets the [[

]]

For LOCA analysis, NuScale assumed that actual core power is 102 percent of rated the core power to account for uncertainty in measured power. The NPM core model assumes that the axial and radial power distribution is at the maximum limits set for core operation.

The NRC staff agrees that the assumption of 102 percent core power along with a power distribution set to the maximum allowed for plant operation reasonably assures that the maximum core operating power prior to a LOCA is conservatively modeled, which complies with the requirements of 10 CFR Part 50, Appendix K.

NuScale models the reactor pressure and CNV metal components as passive heat structures with an appropriate distribution of mass and heat transfer surfaces.

The NRC staff reviewed this basic modelling technique and finds it is acceptable because both RCS pressure vessel metal mass is conservatively modeled, and the conduction heat transfer is appropriately captured.

The NPM reactor model accounts for fission power due to prompt and delayed neutrons using a point kinetics approach. The model simulates reactivity changes due to reactor trip, fuel temperature changes (Doppler coefficient), and moderation changes (moderator density and temperature). The NuScale model sets the moderator and Doppler coefficients for minimum negative worth-based beginning of cycle (BOC) burnup and the reactivity feedback conditions that evolve during the LOCA. The scram rod worth is appropriately delayed for trip and insertion times and accounts for the most reactive rod stuck outside the core.

The point kinetics core model divides delayed neutrons into six precursors groups, and calculates core decay heat in accordance with the 1973 American Nuclear Society (ANS) standard which the NRC staff reviewed and determined was consistent with the 1971 standard that complies with the regulatory requirements of 10 CFR 50.46, Appendix K. The NRC staff determined that this modeling approach meets the regulatory requirements for LOCA analysis.

Pressurizer

The NRC staff reviewed the description of the modeling of the Pressurizer and finds that the Pressurizer model is acceptable because the model properly accounts for the fluid volume, elevation changes and the metal stored energy. It also models the connection of three RVVS, the pressurizer baffle plate, the pressurizer heaters, and the steam plenums interface.

17 Helical Coil Steam Generators (HCSG)

Section 5.1.3, “Helical Coil Steam Generators,” of the NuScale TR describes how NuScale models the two HCSGs within the NPM that is wrapped within the entire cylindrical cross- section of the upper cold side of the RCS loop using a newly developed NRELAP5 SG component model to capture the performance of this unique SG design. The NPM model includes noding required to calculate the heat transfer from the reactor coolant into the HCSG tubes. The HCSG model includes equivalent noding for the secondary coolant inside the HCSG tubes. The flow of the two-phase secondary coolant travels from the lower feedwater supply, into tubes to boil, and then up to the steam headers where it exits as superheated steam, is modeled.

During a LOCA, the SGs are isolated after reactor trip from the remainder of the secondary system but remain active via the DHRS which receives steam from the generators and transfers heat to the pool region, outside of the CNV, by condensation and returns the condensate to the SG. This would provide an additional means of removing decay heating from the core and controlling the core conditions. However, NuScale has chosen not to take credit for the DHRS in the LOCA EM. Thus, for the analysis of LOCAs and the inadvertent opening of ECCS valves, the SGs are isolated with stagnant tubes where heat transfer is then limited to the downcomer region only.

Although the energy transferred to or from the SGs is accounted for in the NRELAP5 LOCA evaluations, the initial operating temperature of the SGs is still high. The most significant factor for the HCSG, is the resistance to flow during early portion of the LOCA, as this influences the heat removal capability in the core. The HCSG secondary side impact on a LOCA, is the heat transfer capability. NuScale has tested both of these factors in prototype testing and incorporated the results of the testing into the new NRELAP5 SG component. The NRC staff notes that the primary impact of the HCSG is in the early phase of a LOCA. The detailed review of the NuScale Società Informazioni Esperienze Termoidrauliche (SIET) TF1 test validation is documented in the NRC staff’s SER of the Non-LOCA EM TR (ML20042E039).

The NRC staff finds that the NuScale approach to HCSG modelling is conservative for both initial steady state and transient analysis. The NRC staff agreed that since the role of the HCSGs are minor in the LOCA transients, the effects of tube-plugging and fouling are also negligible.

Containment Vessel and Reactor Pool

Section 5.1.4, “Containment Vessel and Reactor Pool,” of the NuScale TR describes how the NuScale LOCA model represents the CNV with [[

18 ]]

Chemical and Volume Control System

The LOCA TR describes the modeling of the CVCS in Section 5.1.5, “Chemical and Volume Control System.” Within the CNV, the CVCS system is comprised of small pipes connected to the RPV riser section for supply and the RPV downcomer section for letdown. NuScale stated that [[

]]

The NRC staff finds that the NuScale break spectrum conservatively bounds any LOCA associated with the CVCS due to the consideration of LOCAs in the CVCS piping, the CVCS isolation function, and the fact the model neglects water injected prior to isolation, and therefore, accepts this model for the CVCS.

Secondary System

The LOCA TR describes the modeling of the secondary system in Section 5.1.6. NuScale stated that the SG secondary side is [[

19 ]]

The NRC staff finds that the NuScale LOCA model treatment of the secondary system is acceptable for LOCA evaluations because it treated secondary side energy contribution conservatively.

Decay Heat Removal System

The LOCA TR describes the modeling of the DHRS in Section 5.1.7. The NuScale DHRS design includes Isolation valves, closed during normal operation, that open upon activation of the DHRS. When these DHRS isolation valves open, steam generated in the HCSG tubes by heat transfer from the RCS is condensed in the DHRS heat exchanger by condensation on tubes cooled by ultimate heat sink pool.

For its LOCA EM, NuScale conservatively takes no credit for the DHRS. To justify the NuScale assumption that operation of the DHRS would not have an adverse impact on LOCA results, NuScale performed a break spectrum analysis for which the DHRS was active. The NuScale analysis of the active and inactive DHRS is presented in Section 9.3, “Decay Heat Removal System Availability,” of the LOCA EM TR. The results show that the collapsed water level is not significantly impacted by assuming that the DHRS is active for larger breaks. For smaller breaks, operation of the DHRS prevents re-pressurization of the NPM. The NRC staff finds that the NuScale assumption that the DHRS is inactive, is a conservative assumption for the LOCA EM and is therefore, acceptable.

NRELAP5 Modeling Options

The NuScale NPM LOCA analysis is performed with Version 1.4 of NRELAP5. NRELAP5 uses the ‘h2o95’ water property table for all systems of the LOCA model. The NRELAP5 code has a default feature of termination of the transient if a system mass error exceeds one percent. NRELAP5 uses air as the only non-condensable gas for the partially evacuated CNV and NRELAP5 uses air at normal air pressure above the reactor pool water surface. Because these water property tables and air property are part of previously approved RELAP5 code features, the NRC staff finds these assumptions to be acceptable.

JUNCTION OPTIONS NuScale provided a list of the junction options selected for its LOCA EM in Table 5-1, “Default junction options for the NRELAP5 loss-of-coolant accident model,” of the LOCA EM TR. NuScale used [[ ]]

The NRC staff finds that the NuScale junction options selection options as shown in Table 5-1, are acceptable because they properly model the fluid flow area and local loss coefficients and have added Moody critical flow models, which comply with the regulatory requirements in 10 CFR Part 50, Appendix K to model choke flow.

VOLUME OPTIONS NuScale documented its selection of the volume options for its LOCA EM, in LOCA EM TR Section 5.1.8.2, “Volume Options.” The NRC staff finds that these modeling options are acceptable for this application because they properly model the fluid interphase friction and wall 20 friction consistent with the applicable regulatory requirements in 10 CFR Part 50, Appendix K to model break flow phenomena and friction pressure drops.

HEAT STRUCTURE OPTIONS NuScale discussed its selection of the heat structure options for its LOCA EM, in LOCA EM TR Section 5.1.8.3, “Heat Structure Options.”

The NRC staff notes that NuScale has added several boundary condition types to model unique aspects of the NPM and the NRC staff finds that this heat structure treatment is acceptable for the NuScale LOCA EM because it appropriately identifies that the options related to the FOM being CHF and CLL. In addition, the heat structure components modeled conservatively, accounts for the metal mass, the sensible heat and heat conduction during a LOCA transient.

The NRC staff notes that NRELAP5 v1.4 includes Option 170 with use of the [[ ]] for fuel rod CHF. However, the LOCA EM specifies use of Option 171 with use of the [[ ]]. As such, the NRC staff did not review and does not approve of use of Option 170. Time Step Size Control

NuScale discussed NRELAP5 time step control in LOCA EM TR Section 5.1.9, “Time Step Size Control.” NuScale stated that a sensitivity study was performed to demonstrate that the selected maximum time-step size has no significant impact on the LOCA FOM such as peak containment pressure and CLL in the RPV riser. The NRC staff audited this sensitivity study, as described in the associated audit report (ML20010D112), and finds that the analysis results are not sensitive to the time step size range. Therefore, the NRELAP5 time step selection process using [[ ]] is reasonable.

4.5.2 Analysis Setpoints and Trips

Section 5.2, “Analysis Setpoints and Trips,” of the LOCA EM TR discusses and lists the system trips that NuScale incorporated in the LOCA EM. NuScale stated that signals that are not credited either do not play a role in a LOCA or provide conservatism by delaying actuation of safety-related systems that only reduce the consequences of a LOCA.

The NRC staff finds that the NuScale approach for selecting and modeling analysis setpoints and trips in their LOCA EM is acceptable based on the NRC staff’s review of the trip set points included in the TR, which are the ones that are necessary for appropriately modeling the LOCA events in a conservative manner. Evaluation of the instrumentation is performed under the Design Certification review and is not a subject of this SER.

After the TR was developed, a design change added a low RCS pressure ECCS actuation and removed an RPV level actuation. These changes were not incorporated into the model used in the examples in this TR. However, the staff determined that these changes to the ECCS setpoints would not significantly impact the example calculations used to demonstrate an acceptable EM. Additionally, limitation and condition 5 in Section 6 of this report requires the use of NPM model Revision 3 which includes the updated ECCS setpoint logic .

21 4.5.3 Initial Plant Conditions

In Table 5-6, “Plant initial conditions,” of the LOCA TR, NuScale listed initial plant conditions, including core power, RCS temperature and pressure, pressurizer level, CNV pressure, secondary system pressures and temperatures, and the initial level and temperature for the reactor pool (ultimate heat sink). NuScale states that these initial plant conditions are conservatively biased for LOCA analysis and that the plant conditions are selected to account for both the normal control system deadband and the system/sensor measurement uncertainty.

Section 5.3, “Initial Plant Conditions,” of the LOCA TR lists the process parameters associated with the plant initial conditions, which serves as input the LOCA EM. The NRC staff reviewed NuScale’s LOCA calculations (ML17066A463) to determine whether these parameters were chosen conservatively. Further, the NRC staff performed its own sensitivity studies using NuScale’s NRELAP5 code and input model and varied parameters such as initial pool temperature, RCS temperature and pressure, and pressurizer level. The NRC staff confirmed through its sensitivity studies that the initial plant conditions listed in Table 5-6 are chosen conservatively. For example, [[

]] is conservatively used for the LOCA analysis.

The NRC staff further finds that the NuScale has provided sufficient detail to ensure that the appropriate bounding plant conditions have been selected for each LOCA analysis and that the system and measurement uncertainties established by NuScale in the DCA have been conservatively included in the LOCA analyses.

4.5.4 Loss-of-Coolant Accident Break Spectrum

In Section 5.4, NuScale presents its break location, configuration and size, single failure, loss- of-power, and Decay Heat Removal System availability assumptions as part of its break spectrum definition. The considered break locations are a maximum of 2-inch piping.

Break Location, Configuration and Size

NuScale postulates break locations in the NPM RCS injection and discharge line, pressurizer spray supply line, and high point vent lines. NuScale does not consider the RRV and RVV valve flange connections as break locations. The NRC staff reviewed the locations selected and compared it to the potential break locations and finds this identification of break locations acceptable due to break exclusion. The NRC staff notes that the RRV and RVV break exclusion zone review is performed as part of the Design Certification review in Chapter 3, “Design of Structures, Systems, Components and Equipment.”

The NuScale break spectrum is based on piping that penetrates the RPV wall, connects to the CNV or passes through the CNV. There are four such entities; [[

22 ]] The NRC staff reviewed the break spectrum flow area selection against the design and based on that review, agrees that this selection is appropriate. The NRC staff further noted that the break spectrum demonstrates that the 5 percent injection line break without DHRS operation, loss of alternating current (ac) power, and failure of one ECCS division is the most limiting break case that produces the minimum level. The 100 percent CVCS line break produces the MCHFR. Single Failures

As noted by NuScale in TR Section 5.4.3, “Single Failures,” 10 CFR Part 50, Appendix K requires that single failures be considered within the break spectrum. 10 CFR Part 50, Appendix K requires analyzing single failures of a system or component classified as non-safety related if the inclusion of that system or component would introduce a more limiting condition for LOCA analysis. For single failures, NuScale considered failures of a single RVV or RRV valve to open, or failure of one division of ECCS valves to either actuate or inadvertently actuate.

A single failure inadvertent actuation of a division when it should not be activated, means that direct current (dc) power is removed from that division and two RVVs and one RRV will be available to open, and will do so once the differential pressure (dp) between the RPV and CNV drops below the IAB release pressure. If dc power remains available to the other division, that division’s valves will reposition on a valid actuation signal. If that actuation signal is received later than the IAB release pressure being achieved, then this creates a staggered release of the five ECCS valves (three earlier at the IAB release pressure and two later on the level actuation signal). The applicant stated that this scenario is non-limiting. [[

]] The IAB valve is a first-of-a-kind, safety-significant, active component integral to the NuScale ECCS. To meet the requirements for the ECCS in 10 CFR Part 50, an applicant must show that it has evaluated the single failure criterion (SFC). The SFC is defined in 10 CFR Part 50, Appendix K and derived from the definition of single failure in 10 CFR Part 50, Appendix A. During its review, the NRC staff noted that although the applicant assumed a single failure of a main ECCS valve to open, the applicant did not apply the SFC to the IAB valve regarding the valve’s function to close. NuScale disagreed with the NRC staff’s application of the SFC to the IAB valve, which led the NRC staff to request the Commission’s direction to resolve this issue, SECY-19-0036, “Application of the Single Failure Criterion to NuScale Power LLC’s Inadvertent

23 Actuation Block Valves.”1 In SECY-19-0036, the NRC staff summarized the NRC’s historical practice for applying the SFC. Specifically, the NRC staff summarized SECY-77-439,2 in which it informed the Commission of how the NRC staff then generally applied the SFC, and, SECY- 94-084,3 in which the NRC staff requested the Commission’s direction on the application of the SFC in specified fact- or application-specific circumstances. In view of this historical practice, the NRC staff in SECY-19-0036, requested the Commission’s direction on the application of the SFC to the IAB valve’s function to close. In response to the paper, the Commission directed the NRC staff in SRM-SECY-19-0036, “Staff Requirements - SECY-19-0036 - Application of the Single Failure Criterion to NuScale Power LLC’s Inadvertent Actuation Block Valves,”4 to “review Chapter 15 of the NuScale Design Certification Application without assuming a single active failure of the inadvertent actuation block valve to close.” The Commission further stated that “[t]his approach is consistent with the Commission’s safety goal policy and associated core damage and large release frequency goals and existing Commission direction on the use of risk-informed decision-making, as articulated in the 1995 Policy Statement on the Use of Probabilistic Risk Assessment Methods in Nuclear Regulatory Activities and the White Paper on Risk-Informed and Performance-Based Regulation (in SRM SECY- 98-0144 and Yellow Announcement 99-019).” Based on the NRC staff’s historic application of the SFC and the Commission’s direction on the subject, as described in SECY-77-439, SRM-SECY-94-084, and SRM-SECY-19-0036, the NRC has retained some discretion, fact- or application-specific circumstances, to decide when to apply the SFC. The Commission’s decision in SRM-SECY-19-0036, provides direction regarding the appropriate application and interpretation of the regulatory requirements in 10 CFR Part 50, to the NuScale IAB valve’s function to close. This decision is similar to those documented in previous Commission documents that evaluated the use of the SFC and provided clarification on when to apply the SFC in other specific instances. Specific LOCA event limiting single failures are evaluated as part of a design-specific application of this methodology, such as the NuScale DCA. This is reflected in item 6, in Section 6.0 of this SER.

1 See SECY-19-0036, “Application of the Single Failure Criterion to NuScale Power LLC’s Inadvertent Actuation Block Valves,” (April 11, 2019) (ADAMS Accession No. ML19060A081).

2 See SECY-77-439, "Single Failure Criterion," (August 17, 1977) (ADAMS Accession No. ML060260236).

3 SECY-94-084, "Policy and Technical Issues Associated with the Regulatory Treatment of Non-Safety Systems in Passive Plant Designs (March 28, 1994) (ADAMS Accession No. ML003708068), and associated SRM (June 30, 1994) (ADAMS Accession No. ML003708098).

4 See SRM-SECY-19-0036, “SECY-19-0036 Application of the Single Failure Criterion to NuScale Power LLC’s Inadvertent Actuation Block Valves,” (July 2, 2019) (ADAMS Accession No. ML19183A408). 24 Loss-of-Power

The NuScale LOCA evaluation methodology considers two scenarios for loss-of-power coincident with a postulated LOCA:

• Complete loss of normal ac and dc power • Complete loss of only ac power with dc power availability

The NRC staff finds that the NuScale LOCA EM appropriately models the impacts of loss of ac and/or dc power coincident with a LOCA because it considers both scenarios.

Specific LOCA event limiting electric power assumptions are evaluated as part of a design- specific application of this methodology, such as the NuScale DCA. This is reflected in item 6 in Section 6.0 of this SER. Decay Heat Removal System Availability

The NuScale LOCA EM does not credit the function of the DHRS. However, in Section 9.3, “Decay Heat Removal System Availability,” of the EM TR, NuScale documented a sensitivity study that shows the beneficial outcome of the DHRS operation that the NRC staff reviewed. Because the DHRS operation is beneficial during a LOCA, the NRC staff finds that the NuScale assumption that the DHRS does not function, is conservative for the NuScale LOCA EM because it does not include this potential benefit from DHRS operation.

4.5.5 Sensitivity Studies

For the NRC staff’s evaluation of NuScale’s sensitivity studies, please see Section 4.9 of this SER.

4.5.6 Review Focus of TR Section 5

NRC regulation 10 CFR 50.46 requires applicants to show that they have analyzed the bounding break within the break spectrum relative to the FOMs for LOCA evaluations. The most critical FOMs for the NuScale LOCA EM are to show that minimum CLLs in the NPM riser remain above the active core and that the maximum fuel heat flux remains below the CHF. The NRC staff finds that the NuScale LOCA EM is capable of calculating these FOMs. The representative LOCA evaluations included in the NuScale LOCA EM TR, are meant to be examples to demonstrate the EM’s capability but do not necessarily identify the most limiting LOCA, as this will need to be done through application of the methodology to a design. For example, as described in Section 4.2 of the TR, the example calculations were completed prior to the NPM design change to add an ECCS actuation on low RCS pressure. The staff determined that the example calculations presented in the TR reasonably represent the capability of the methodology.

4.5.7 Overall Conclusions of the review of TR Section 5

Subject to the limitations discussed above, the NRC staff finds that the NuScale LOCA EM, as described in LOCA EM TR-0516-49422-P, Section 5, is acceptable for referencing in the NuScale DCA because the modeling developed for the EM is deemed adequate to represent the key phenomena and features of the NPM. 25 4.6 NRELAP5 Computer Code

As stated in the LOCA EM TR, Section 6.0, NuScale used its proprietary NRELAP5 system thermal-hydraulics code for evaluating small break LOCA ECCS performance. This NuScale NRELAP5 code, developed from RELAP5-3D©, Version 4.1.3, includes hydrodynamics, heat transfer between structures and fluids, modeling of fuel, reactor kinetics models, and control systems models. Like previous versions of RELAP5 codes, the NuScale NRELAP5 code used a two-fluid, non-equilibrium, non-homogenous model to simulate system thermal-hydraulic response. In Section 6.0, NuScale provided a general overview of the NRELAP5 computer code structure, models, and correlations and a description of the LOCA code models and code changes implemented by NuScale to model unique design features and phenomena for the NPM.

NuScale added or revised the following models to NRELAP5, following the requirements of the NuScale QAP:

• [[ • • • • ]]

The NRC staff’s review of the NuScale NRELAP5 computer code focused only on NuScale changes and additions to the RELAP5-3D© code, and, the applicability of the NRELAP5 code to the NPM. The NRC staff’s review was based on NRELAP5 Version 1.4. This is reflected in item 5 in Section 6.0 of this SER.

4.6.1 Quality Assurance Requirements

Compliance with QA requirements is described in “NuScale Topical Report: Quality Assurance Program Description for the NuScale Power Plant,” NP-TR-1010-859-NP-A. The NRC staff reviewed the Quality Assurance Program Description and documented its approval in its SER (ML16347A405). Further, the NRC staff inspected NuScale’s design control process and code development procedures. These inspections are documented in inspection reports dated October 7, 2017 (ML15268A186) and July 24, 2017 (ML17201J382).

4.6.2 NRELAP5 Hydrodynamic Model

In TR Section 6.2, “NRELAP5 Hydrodynamic Model,” NuScale stated that the NRELAP5 hydrodynamic model is a transient, two-fluid model for flow of a two-phase vapor/gas-liquid mixture that can contain non-condensable components in the vapor/gas phase as well as a soluble component (i.e., boron) in the liquid phase. The NRELAP5 two-fluid equations of motion are formulated in terms of volume and time-averaged parameters of the flow. For most analyses, NRELAP5 uses empirical formulas to calculate bulk properties, such as friction and heat transfer.

Because the NRELAP5 hydrodynamic model framework is essentially identical to the RELAP5- 3D© code and the NRC staff has approved its application for LOCA analyses of the US EPR 26 design (ML110070113) and the US APR1400 design (ML18180A327), the NRC staff only reviewed the code changes and their applicability of the unique aspects to the NPM.

Field Equations

In Section 6.2.1, “Field Equations,” of the LOCA TR, NuScale discusses the NRELAP5 thermal- hydraulic model. Because NRELAP5 basic field equations 6-1 to 6-10 are the same as the field equations in the RELAP5-3D© code (and the same equations as contained in the RELAP5/Mod3 code which has been thoroughly reviewed in the past by the NRC staff, as submitted by other vendors for review and approval), the NRC staff did not perform an in-depth review of these field equations and numerical solution techniques. The NRC staff confirmed that NRELAP5 correctly addressed NRC Information Notices 92-02 and 92-02, Supplement 1. Specifically, NuScale demonstrated that when the RELAP5 code is applied to situations in which the pressure drops significantly between cells, the energy in the downstream volume is not significantly underestimated), as INL modified RELAP5 to successfully resolve this issue. The modification preserves the total enthalpy across a junction where a large pressure gradient exists, as in a blow down event into a CNV, for example, where dissipation terms become more important. This modification also resolved the issues raised by NRC Information Notices 92-02 and 92-02, Supplement 1, as the modification is shown to conserve energy transfer across a junction where a large differential pressure exists. The NRC staff reviewed the break and/or orifice input model junction flags to ensure that they are appropriately set to involve the energy correction modeling (e=1).

The NRC staff audited the NRELAP5 Theory manual regarding code modifications made to the NRELAP5 code for application to the EM with focus on the changes from the INL RELAP5-3D code Version v.4.1.3, as described in the associated audit reports (ML20010D112 and ML20034D464).

As discussed in the associated audit report, the NRC staff audited NuScale’s comparison of the NRELAP5 code prediction with known solutions for a simple oscillating manomenter, which showed that the NRELAP5 code properly predicts level and flow behavior for this benchmark exercise. This confirms that the treatment of the inertia term does not introduce significant error into the ability of the code to capture the correct amplitude and period of the oscillations. Adequate prediction of a simple manometer provides validation that NRELAP5 properly models hydrostatic forces that govern small break LOCA behavior as well as to assure the NRC staff that the momentum equation, with and without friction, is properly formulated and implemented in the code.

To confirm that the NRELAP5 code calculations do not show non-physical flow anomalies that would impact analysis of small break LOCAs, the NRC staff audited NuScale’s results of the prediction of fluid flow behavior in a simple system containing parallel pipe components, as described in the associated audit report (ML20010D112) The NRC staff noted that flow anomalies that had been present in earlier versions of RELAP5, were not present in the NRELAP5 version of RELAP5-3D. The core nodalization consists of the application of the 1-D modeling technique in NRELAP5. The NRC staff further recognizes that 3-D modeling of a core is not necessary to accurately predict two-phase level swell following a small break LOCA. The 1-D and 3-D predictions of small break LOCA two-phase level swell have been shown to be in very close agreement, as the multi-dimensional flow capability does not cause the two-phase level to vary significantly across the radius of the nuclear core. NuScale further proved that the different nodalization included in NRELAP5 resolves the anomalous fluid behavior as 27

exemplified by comparison to the dual and triple parallel pipe problems. Therefore, the NRC staff concluded that the current NRELAP5 core modeling would avoid the potential flow anomalies and that the 1-D channel model is reasonably accurate based on additional information provided by NuScale (ML18031B319).

State Relations

In the LOCA TR Section 6.2.2, NuScale discusses the NRELAP5 six-equation model based on five independent state variables with an additional equation for the non-condensable gas component. Because these state equations are the same as the field equations in the RELAP5- 3D© code and its predecessors, which were reviewed and approved before, these state equations are acceptable.

Flow Regime Maps

In Section 6.2.3, “Flow Regime Maps,” NuScale stated that one-dimensional field equations for the two-fluid model used in NRELAP5 precludes direct calculation of physical parameters, such as velocity or energy, that depend upon transverse gradients. Therefore, NRELAP5 adds algebraic terms to the conservation equations for a specific flow regime to provide closure to the two-fluid equations. NRELAP5 flow regime maps are based on the work of Taitel and Dukier and Ishii, as referenced in Section 6.2.3 of the LOCA TR, but further simplified by NuScale to efficiently apply these criteria in NRELAP5. A schematic of the vertical flow regime map, as coded in NRELAP5, is shown in LOCA TR Figure 6-1, “Schematic of Vertical flow-regime map indicating transitions,” to illustrate flow-regime transitions as functions of void fraction, average mixture velocity and boiling. The NRELAP5 junction map is shown in LOCA TR Section 6.2.3.2, “Junction Flow Regime Maps.” The NRELAP5 flow regime maps used for junctions are the same as used for the volumes and are based on the work of Taitel and Dukler, Ishii and Tandon, as referenced in Section 6.2.3 of the LOCA TR.

Because the NRELAP5 flow regime maps are the same as those in the RELAP5-3D© code and its predecessors, which were reviewed and approved before, the NRC staff considers these flow regime maps to be applicable to the NuScale application.

Momentum Closure Relations

In LOCA TR Section 6.2.4, “Momentum Closure Relations,” NuScale states that NRELAP5 uses two different models for the phasic interfacial friction force computation: the drift flux method and the drag coefficient method. These are same models used in the base version RELAP5-3D© except for the revisions NuScale made to implement the new HCSG component.

NRELAP5 uses the drift flux approach only for bubbly and slug-flow regimes for vertical flow. The NRELAP5 drift flux equations are shown in TR Section 6.2.4. NRELAP5 uses the drag coefficient approach in all flow regimes other than vertical bubbly and slug-flow, as described in the equations in Section 6.2.4 of the TR. NRELAP5 determines wall friction based on the volume flow regime map. Because the NRELAP5 momentum equations in Section 6.2.4 of the LOCA TR are the same as the equations in the RELAP5-3D© code and its predecessors, which were reviewed and approved before, these flow regime maps are applicable to the NuScale application.

Heat Transfer

Section 6.2.5, “Heat Transfer,” of the LOCA TR describes the heat transfer equations. NRELAP5 solves for liquid and vapor/gas energy including energy added or removed by the heat flux to or from wall heat structures. NRELAP5 uses boiling heat transfer correlations when the wall surface temperature is above the saturation temperature. When a hydraulic volume is voided and the adjacent surface temperature is subcooled, vapor condensation on the surface is predicted. If non-condensable gases are present, the phenomena are more complex because condensation is based on the partial pressure of vapors present in the region. When the wall temperature is less than the saturation temperature based on total pressure, but greater than the saturation temperature based on vapor partial pressure, a convection condition exists. LOCA TR Figure 6-2, “NRELAP5 boiling and condensing curves,” illustrates these three regions of the curve.

NRELAP5 uses the Chen boiling correlation up to the CHF point. NRELAP5 will issue a message and stop running if CHFR reduces below one for core heat transfer. If the CHFR is below one on other structures outside of the core, NRELAP5 will calculate Post-CHF heat transfer on these surfaces outside the core. NuScale added this stop function to NRELAP5 when the core CHFR drops below one because maintaining core CHFR above one is a critical FOM for the NuScale LOCA EM.

The NRC staff finds that the addition of the stop function to NRELAP5 is appropriate because acceptability of the NuScale LOCA EM depends on maintaining the heat flux from the reactor fuel to the RCS below the CHF point. The detailed review of CHF correlations is documented in Section 5 of this SER.

4.6.3 Heat Structure Models

As discussed in Section 6.3, Heat Structure Models,” NRELAP5 calculates heat transfer from hydrodynamic volumes to adjacent solid heat structures. NRELAP5 has the capability to model various heat structures, allows the user to use standard thermal conductivities and heat capacities or input tables or functions and solves the one-dimensional heat equation with a finite difference method. NRELAP5 allows the user to specify spacing, internal heat source and material composition for each mesh. For nuclear fuel, NRELAP5 calculates the heat source with a reactor kinetics model, or tables of power versus time, or a control system variable. NRELAP5 also includes options for boundary conditions. These modeling features are typical for light water reactor applications. Therefore, they are applicable to NuScale LOCA analyses.

LOCA TR Section 6.3 also describes that the NRELAP5 has heat transfer correlations and a gap conduction model. NRELAP5 solves the heat conduction equation using the Crank- Nicolson method referenced in Section 6.3 of the LOCA TR.

Because the NRELAP5 heat structure and heat transfer models and equations discussed above are the same as those in the RELAP5-3D© code, the NRC staff did not perform an in-depth review of these heat structure and heat transfer models. However, as discussed below, the NRC staff did perform in-depth reviews of the specific heat structure modeling added to NRELAP5, including modeling of steam condensation on the inside wall of the CNV and heat transfer for the HCSG.

4.6.4 Point Reactor Kinetics Model

As described in LOCA TR Section 6.4, “Point Reactor Kinetics Model,” NRELAP5 calculates the total reactor core power from a user specified table or with a point-reactor kinetics model with reactivity feedback. The model uses the ANS 1973 decay heat standard to calculate reactor core power from decay of fission products. The NRC determined that this is similar to ANS 1971, but with higher accuracy, and in compliance with 10 CFR Part 50, Appendix K. Therefore, the NRC staff considers this modeling acceptable.

Furthermore, the selection of the delayed neutron fraction for the kinetics calculation can be justified as conservative for the core in the as-used state. This is typically done as part of the neutronics analysis of the core for a specific cycle design.

The staff confirmed that the NRELAP5 reactor core power and fission power models and equations discussed above are the same as those in the RELAP5-3D© code and its predecessors, which have been reviewed and approved before. The NRC staff found that the point kinetics modeling used is adequate to conservatively calculate the fission power and the decay heat power during a LOCA transient.

4.6.5 Trips and Control System Models

The NRELAP5 modelling of trip and control systems is described in Section 6.5, “Trips and Control System Models.” NRELAP5 provides several types of control variables based on NRELAP5 calculated parameters for each hydrodynamic volume, junction, pump, valve, heat structure, and reactor kinetics. Because the NRELAP5 trip and control system models are the same as those in the RELAP5-3D© code and its predecessors, which were reviewed and approved before, these NRELAP5 trip and control system models are acceptable for NuScale applications.

4.6.6 Special Solution Techniques

As stated in Section 6.6, “Special Solution Techniques,” NRELAP5 uses empirical models for certain processes that are too complex for the general solutions provided in NRELAP5. The NRC staff’s evaluation of these special NRELAP5 empirical models are discussed below.

Choked Flow

MOODY CRITICAL FLOW MODEL NRELAP5 uses the Moody critical flow model, when the break flow is calculated to be two- phase, to comply with the 10 CFR Part 50, Appendix K requirements. NRELAP5 includes options, as described in Section 6.6.1, “Choked Flow,” for switching between [[

]]. Abrupt Area Change

As discussed by NuScale in Section 6.6.2, “Abrupt Area Change,” the NRELAP5 hydrodynamic model provides analytical models for sudden area changes and orifices. NRELAP5 models abrupt area changes with the Borda-Carnot formulation for a sudden enlargement and the vena- contracta effect for a sudden contraction or sharp-edge orifice or both. This formulation does not include models for rounded or beveled enlargements, contractions, or orifices.

Because the NRELAP5 abrupt area change models are the same as those in the RELAP5-3D© code and its predecessor codes, which were approved for LOCA analyses before, the NRC staff considers these models to be acceptable.

Counter Current Flow Limitation

NRELAP5 implements the general CCFL model in a form proposed by Bankoff which has the structure shown in TR Section 6.6.3, “Counter Current Flow Limitation.” NuScale provided an assessment of the NRELAP5 CCFL model against the Bankoff perforated plate test data in LOCA EM TR Section 7.2.10, “Bankoff Perforated Plate,” and NuScale presented a sensitivity study of the effects of the CCFL as it applies to the NPM pressurizer baffle plate in TR Section 9.6.3 “Counter Current Flow Limitation Behaviour on Pressurizer Baffle Plate.”

Because the NRELAP5 uses essentially the same CCFL model as the RELAP5-3D© code, the NRC staff did not perform an in-depth review of these NRELAP5 code implementation of the Bankoff CCFL model in NRELAP5. However, the NRC staff reviewed the assessment of the Bankoff model versus test data as shown in SER Section 4.7.2, “Legacy Test Data,” and the NRC staff evaluated the NuScale Bankoff sensitivity study as shown in SER Section 4.9.7, “Sensitivity Studies.” Since the counter current flow in the NuScale reactor is not a dominant physical phenomenon when the water level reaches the minimum value, this model is acceptable.

4.6.7 Helical Coil Steam Generator Component

As described in Section 6.7, “Helical Coil Steam Generator Component,” NuScale added a new hydrodynamic component and heat transfer package to the NRELAP5 code to model flow and heat transfer inside and outside the HCSG tubes. These added models are specific to helical coil geometry heat transfer and wall friction correlations and were added because the models in the baseline RELAP5-3D© code did not provide adequate agreement with pressure drop and heat transfer performance against prototypic HCSG testing performed at SIET.

The adequacy of the added NRELAP5 HCSG were demonstrated by NuScale through prototypic assessments of the NuScale HCSG using SIET test data. These tests assessed heat transfer and pressure drop on both the secondary side (within tubes) and primary side (external to tubes) of the HCSG that showed good agreement with HCSG tube axial wall and secondary fluid temperature data.

31 The analysis of a LOCA depends on the initial stored energy in the primary coolant and the performance of the NPM HCSG can influence the temperatures and flow rates in the RPV. As described in the associated audit report (ML20010D112), the NRC staff audited a sensitivity study for a variation in the heat transfer performance of the HCSGs, above and below the nominal expected performance. The NRC staff also considered in its audit, the potential distortion on core inlet temperature and SG steam temperature relative to the uncertainty for NIST-1 test results.

The analyses audited included the effect of SG degradation on the initial conditions as well as a typical LOCA progression. Because the NRELAP5 LOCA model conservatively assumes no plugging or fouling in the HCSGs and no credit for DHRS cooling, the SG model was reviewed in detail in the NRC staff’s SER for the NuScale Non-LOCA EM TR (ML20042E039). [[

]] Helical Coil Tube Friction

NuScale implemented SG tube friction models into NRELAP5 for single phase and two-phase flow conditions. [[

]] Therefore, the NRC staff considers the in-tube friction model acceptable for LOCA analyses. Helical Coil Tube Heat Transfer

A new heat transfer package has also been added to NRELAP5 and differs from that of the standard RELAP5 pipe geometry in [[

]]

The transition to [[ ]]

The laminar heat transfer correlation developed by [[

]]

The NRC staff finds the friction and heat transfer models to be acceptable for the LOCA evaluations based on the good agreement with the SIET separate effects heated tube wall and pressure drop data. Further detailed evaluation is documented in the NRC staff’s SER of the Non-LOCA EM TR (ML20042E039).

32 4.6.8 Wall Condensation

As discussed in Section 6.8, “Wall Heat Transfer and Condensation,” the RELAP5-3D© code includes [[

]]

The NRC staff reviewed and audited the NRELAP5 code, as discussed in the associated audit report (ML20010D112), and noted that [[ ]] Since condensation is of major significance to the prediction of ECCS performance, the NRC staff reviewed the information provided by NuScale (ML17324B392 and ML19240C658) regarding the modeling of heat transfer from the CNV to the pool for a loss of primary coolant from the RPV and audited additional calculations underlying the submitted information, as described in the associated audit report (ML20010D112). The NRC staff’s review of use of the extended Shah correlation to model heat transfer in the DHRS was not reviewed as part of this TR, since DHRS heat transfer is not credited. The NRC staff’s review of its use to model heat transfer in the DHRS is documented in the NRC staff’s SER of the Non- LOCA EM TR (ML20042E039).

The information that the NRC staff reviewed, indicated that during [[

]] the NRC staff conducted an audit, as described in the associated audit report (ML19282C504), to determine [[

33 .]] As described in the associated audit report (ML19282C504), the NRC staff also audited [[

34 ]]

The NRC staff finds that these modeling approaches are conservative for calculating minimum CNV heat transfer for maximum peak containment pressure and minimum collapsed water level above the core.

4.6.9 Interfacial Drag in Large Diameter Pipes

[[ ]]

As the NRELAP5 code assessment against General Electric (GE) level swell test showed reasonable agreement between the measured and the calculated as shown in Section 7.2.2, “GE Level Swell (1 ft),” the interfacial drag model for large diameters is reasonably accurate. Thus, this is applicable to the NuScale LOCA analyses.

4.6.10 Fission Decay Heat and Actinide Models

The NRELAP5 implementation of the ANS 1973 standard applies the Shure curve. Comparison of the ANS 1973 standard to the as implemented curve in NRELAP5 shows that the implemented curve reproduces the 1973 standard decay heat data.

The implemented model yields the result quoted in the 1979 Standard, the 1994 Standard, and the 2005 Standard. The 1973 actinide equations are identical to those in the 1979 standard. Comparison of the NRELAP5 model with this standard shows identical results.

Furthermore, infinite operation is assumed and a decay heat multiplier of 1.2 is employed as required by 10 CFR Part 50, Appendix K. The NRC staff notes that previous studies of the various decay heat standards identified the need to include the contributions from additional actinides (other than 239U and 239Pu), since actinide contribution grows significantly with 35 shutdown time. Because the decay heat model used meets the applicable requirements of 10 CFR Part 50, Appendix K by assuming infinite full power operation and the approved ANS 1973 decay heat curve, the NRC staff considers this model to be acceptable.

4.6.11 Critical Heat Flux Models

The NRC staff’s review of the proposed CHF correlations is documented in Section 5.4 of this SER.

4.7 NRELAP5 Assessments

Section 7, “NRELAP5 Assessments,” provides a summary of the NuScale assessments of the SET and IET that NuScale performed. NuScale discussed the comparison of the NRELAP5 analysis of these separate and integral effects tests versus experimental data in Section 8.0, “Assessment of Evaluation Model Adequacy,” and presents its justification of the adequacy for modeling of the high-ranked phenomena in the NuScale LOCA PIRT.

The NRC staff reviewed the separate and integral effects tests and focused on determining the acceptability of the NuScale LOCA evaluation methodology for performing design basis LOCA analyses. This NRC staff review was limited to the applicability of NuScale methodology and use of the NRELAP5 computer code to perform LOCA analysis for the break spectrum as defined by NuScale.

4.7.1 Assessment Methodology

NuScale used various special and integral experimental tests, and analytic problems to assess the performance of NRELAP5 using the process identified in Element 2 of RG 1.203. NuScale chose the tests and analytical problems to assess the adequacy of the NRELAP5 code to model the high-ranked phenomena shown in the NuScale LOCA PIRT as discussed in Section 4 of the LOCA TR. The NRC staff concludes that this process is consistent with that of RG 1.203 and is therefore, acceptable.

4.7.2 Legacy Test Data

Tests that NuScale evaluated in Section 7 were performed by others and were not done in compliance with the NuScale QAP. With the exception of Marviken JIT-11 data, NuScale qualified these tests by applying non-mandatory guidance provided by NQA-1 2008/2009 Addendum. NuScale used Marviken JIT-11 data based on published literature data. Because these legacy test results have been reviewed by the NRC staff previously for several RELAP5 code-based LOCA EM methods, the use of these data by NuScale is acceptable.

Ferrell-McGee

The Ferrell-McGee tests were performed in vertical pipes over a wide range of single phase and two-phase flow conditions with uniform, contraction, and expansion flow areas. NuScale performed analysis of these tests with NRELAP5 and compared the calculations with the experimental data to assess the ability of NRELAP5 code to calculate single- and two-phase pressure drop and void fraction under different pressures, flow rates, and inlet quality.

36 NuScale’s NRELAP5 calculations and comparison to the Ferrell-McGee, are summarized in the LOCA EM TR. The NRC staff audited the calculations underlying the summary in the LOCA TR, as described in the associated audit report (ML20010D112), and found that the NuScale NRELAP5 code calculations show excellent agreement with test data for pressure drop in the bubbly to slug flow regime and satisfactory agreement with the test data in the annular-mist regime. Ferrell-McGee tests at void fractions approaching 1.0 where not usable for comparison to NRELAP5 analysis because the void fractions near 1.0 cannot be measured with sufficient accuracy and because pressure drop is strongly dependent on void fraction. The NRC staff found that NRELAP5 was able to adequately calculate void distribution for all of the Ferrell- McGee test cases based on the observed agreement between the measured and the calculated void fraction distribution. The difference between NRELAP5 calculations and measured pressure drop decreases with increased flow rate, increased pressure and increased hydraulic diameter.

GE Level Swell Test – 1 ft

NuScale assessed NRELAP5’s ability to predict void distribution and level swell phenomena for depressurization transients by assessing it against the GE Level Swell Test referenced in Section 7.2.2, “GE Level Swell (1 ft).” The specifics of the test configuration and instrumentation are described by NuScale in its LOCA TR and the NRC staff reviewed these descriptions.

The GE Level Swell Test 1004-3 is a small-break blowdown of a vertical vessel for which GE measured the axial void fraction distribution. NuScale modeled the GE test facility and compared the two-fluid interphase level calculated by NRELAP5 to the measured void fraction distributions from the GE test. NuScale used these comparisons to assess whether NRELAP5 correctly predicts single phase and two-phase choked flow, liquid level, flashing, level swell, mixture level and phase slip and flow. NuScale used the Henry Fauske critical flow correlation with four contraction coefficients (1.0, 0.9, 0.7 and 0.6) to calculate the break flow. NuScale provided sensitivity analyses for blowdown line orientation and vessel nodalization.

NuScale determined that the selection of the Henry Fauske critical flow correlation with a 0.9 discharge coefficient provides the best comparison of NRELAP5 calculated vessel pressure to the GE test data. NuScale also determined that NRELAP5 analysis results are not sensitive to the other modeling options. The NuScale NRELAP5 model of the GE 1-foot (ft) (0.3 meters (m)) vessel generally over predicts void fraction. NRELAP5 only under predicts void fraction at the 12-feet (3.7m) elevation for times of 10 and 40 seconds.

Because the primary FOM as shown in TR Section 4 for NuScale LOCA analyses is a CLL in the NPM riser, the NRC staff agrees with NuScale that the NRELAP5 predicted void fractions are in reasonable agreement with the measured data and the Henry-Fauske critical flow correlations should be used for break flow in the subcooled region and Kataoka-Ishii and Zuber- Findlay for the interfacial drag model in pipes as discussed in Section 6.9.

GE Level Swell Test – 4 ft

NuScale assessed NRELAP5’s ability to predict void distribution and level swell phenomena for depressurization transients by assessing it against the GE Level Swell Test referenced in Section 7.2.3, “GE Level Swell (4 ft).” The specifics of the test configuration and

37 instrumentation are described by NuScale in its LOCA TR and the NRC staff reviewed these descriptions.

The GE 4-ft (1.2m) tank level swell tests measured time-dependent pressures and void fraction profiles in a large tank which was depressurized via a blowdown line. NuScale modeled the GE test facility and compared the two-fluid interphase level calculated by NRELAP5 to the measured void fraction distributions from the GE test. NuScale used these comparisons to assess whether NRELAP5 correctly predicts single phase and two-phase choked flow, liquid level, flashing, level swell, mixture level and phase slip and flow.

[[

]]

Because the primary FOM as shown in TR Section 4 for NuScale LOCA analyses is a CLL in the NPM riser, the NRC staff found that the assessment results are in support of using the Henry-Fauske correlation for the critical break flow calculation in the subcooled region and Kataoka-Ishii and Zuber-Findlay for the interfacial drag model in the riser.

KAIST

NuScale assessed the DHRS condensation modeling of its NRELAP5 code against experimental data from KAIST. Since the assessment was relative to the behavior of DHRS, which is not credited in the LOCA analysis, the review results of the assessment analysis are documented in the NRC staff’s SER of the Non-LOCA EM TR (ML20042E039).

FRIGG

NuScale assessed NRELAP5’s ability to model interphase drag and heat transfer models under two phase flow conditions by assessing it against the FRIGG Tests referenced in Section 7.2.5, “FRIGG.” The specifics of the test configuration and instrumentation are described by NuScale in its LOCA TR and the NRC staff reviewed these descriptions.

NuScale modeled the FRIGG test facility and compared the void distribution data to the measured void distributions from the FRIGG tests. NuScale used these comparisons to assess whether NRELAP5 correctly predicts the void distribution data in a rod bundle geometry as a function of mass flow, inlet subcooling, system pressure and thermal power.

Four tests were chosen for assessment. The NRC staff agrees with NuScale’s conclusion that Figures 7-23 to 7-26 of the LOCA TR show that NRELAP5 predicted the experimental void fraction data with reasonable agreement, justifying use of one dimensional nodalization to obtain reasonable predictions of the axial void profile. These results validate the NRELAP5 interphase drag and heat transfer models for applications having similar core geometries.

FLECHT-SEASET

NuScale assessed NRELAP5’s ability to model bundle boil-off by assessing it against the FLECHT-SEASET tests referenced in Section 7.2.6 of the LOCA TR. The specifics of the test

configuration and instrumentation are described by NuScale in its LOCA TR and the NRC staff reviewed these descriptions.

NuScale modeled the FLECHT-SEASET test facility and compared the void distribution data to the measured void distributions from the FLECHT-SEASET tests. NuScale used these comparisons to assess whether NRELAP5 correctly predicts the void distribution data for a core boiloff configuration. The NRC staff audited NuScale’s comparison calculations, as described in the associated audit report (ML20010D112), and noted a significant difference in the early part of the transient. As shown in Figure 7-28, “FLECHT-SEASET level 1 void fraction versus time – Test 35557,” of the LOCA TR, voids appeared almost immediately after the initiation of the transient in the calculations, whereas there was a delay in the void generation in the experimental data. It appears to the NRC staff that the difference between the calculations and the data may be a time delay in the heatup of the rods. The LOCA TR shows the calculated void fraction history at various levels in the test section compared to data for one of the three tests. The calculations represented the trend of the data reasonably well. Early in time and at the lower levels, it appears the calculated entrainment rate is too high and thus the void fraction is over-calculated. The entrained liquid is carried up and out of the test section as evidenced by the lower calculated void fraction at elevations above the bottom cell during that time. This behavior also persists at later times as observed in the figures in LOCA TR Section 7.2.6. Simulation of the boiloff test seems to indicate that the interphase drag calculated by the code is too large. The rate of coolant lost out the bundle top in the calculation is greater than shown by the data. Figure 7-1, “Schematic of the Ferrell-McGee test section,” of the LOCA TR indicates that these tests partially evaluated subcooled boiling at the spacers. Although there is some level of deviation between the measured void fraction and the calculated void fraction, the NRC staff considered that the NPM modeling derived from the FLECHT-SEASET boiloff test is reasonable because there is a large amount of collapsed water level above the core for the NuScale design.

SemiScale (S-NC-02 and S-NC-10)

NuScale assessed NRELAP5’s ability to model [[ ]] by assessing it against the SemiScale tests referenced in Section 7.2.7, “SemiScale (S-NC-02 and S-NC-10),” of the LOCA TR. The specifics of the test configuration and instrumentation are described by NuScale in its LOCA TR and the NRC staff reviewed these descriptions. NuScale modeled the SemiScale test facility and compared the loop mass flowrate as a function of system inventory to the measured data from the SemiScale tests. NuScale used these comparisons to assess whether NRELAP5 correctly predicts natural circulation flow. For the test S-NC-2, the calculated results of flow versus inventory are under predicted at higher inventories and are adequately predicted at lower inventories for both power levels. In the 75- 80 percent inventory for test S-NC-2, the code exhibits an oscillatory behavior. As described in the associated audit report (ML20010D112), the NRC staff audited NuScale’s assessment, which attributed this to flow regime flip-flopping in the lowest core node. For test S-NC-10 at 100 kW, the calculated results compare well with the data in the 97 percent to 100 percent mass inventory range. In the lower inventory range, the flow rates were over predicted by as much as 40 percent. The over prediction is attributed by the applicant to the lowest node having a bubbly flow regime, resulting in more interfacial area and thus more drag compared to a slug regime, which would result in lesser drag force. However, since the assessment focused on the natural circulation and the NRELAP5 prediction matched the measured flow rate well, the code assessment against SemiScale tests are acceptable. 39

Wilson Bubble Rise

NuScale assessed NRELAP5’s ability to model the void fraction distribution in the hot leg riser by assessing it against the Wilson Bubble Rise test referenced in Section 7.2.8, “Wilson Bubble Rise,” of the LOCA TR. The specifics of the test configuration and instrumentation are described by NuScale in its LOCA TR and the NRC staff reviewed these descriptions.

NuScale modeled the Wilson Bubble Rise test facility and compared the void fraction at different pressures to the measured data from the Wilson Bubble Rise test. NuScale used these comparisons to assess whether NRELAP5 correctly predicts void fraction distribution in the hot leg riser. NuScale’s assessment results show the void fraction is over predicted at lower flow rates and under predicted at higher flow rates. However, the overall comparison is within a normal range of error band. Therefore, the NRC staff finds this part of assessment acceptable.

Marviken Jet Impingement Test

NuScale assessed NRELAP5’s single phase choked flow model by assessing it against the Marviken Jet Impingement test referenced in Section 7.2.9, “Marviken Jet Impingement Test (JIT) 11,” of the LOCA TR. The specifics of the test configuration and instrumentation are described by NuScale in its LOCA TR and the NRC staff reviewed these descriptions.

NuScale modeled the Marviken Jet Impingement test facility and compared simulated mass flow rate and density for various values of the discharge coefficient to the measured data from the Marviken Jet Impingement test.

NuScale applied the [[

]]. Therefore, the NRC staff considers the code assessment against the Marviken test acceptable. Bankoff Perforated Plate

NuScale assessed NRELAP5’s ability to model countercurrent flow by assessing it against the Bankoff Perforated Plate test referenced in Section 7.2.10, “Bankoff Perforated Plate,” of the LOCA TR. The specifics of the test configuration and instrumentation are described by NuScale in its LOCA TR and the NRC staff reviewed these descriptions.

NuScale modeled the Bankoff Perforated Plate test facility and compared the vapor superficial velocity to the measured data from the Bankoff Perforated Plate test. NuScale used these comparisons to assess whether NRELAP5 correctly predicts countercurrent flow. The NRC staff reviewed this comparison, which shows that NRELAP5 predictions are in excellent agreement, thus demonstrating that the correlation is correctly implemented in NRELAP5 and that the code can accurately model the countercurrent flow phenomena that occurs in the Bankoff tests.

Marviken Critical Flow Tests 22 and 24

NuScale assessed NRELAP5’s ability to model blowdown conditions where discharge flow is limited by choked conditions by assessing it against the Marviken Critical Flow tests referenced in Section 7.2.11 of the LOCA TR. The specifics of the test configuration and instrumentation are described by NuScale in its LOCA TR and the NRC staff reviewed these descriptions. 40

NuScale modeled the Marviken Critical Flow test facility and compared the mixture density to the measured data from the Marviken Critical Flow tests. NuScale used these comparisons to assess whether NRELAP5 correctly predicts critical flow in piping breaks.

The NRC staff reviewed NuScale’s sensitivity study, including sensitivity to time step, nodalization and critical flow at the break and the results for the four different critical flow correlations tested, as described in Section 7.2.11, “Marviken Critical Flow Test 22 and 21,” of the LOCA TR. The NRC staff concludes that NuScale’s analysis shows that NRELAP5 has the capability to perform critical flow calculations with reasonable agreement to test data.

4.7.3 NuScale Critical Heat Flux Tests

The NRC staff’s review of the proposed CHF correlations, including the test data used in the development and validation of the correlation, is documented in Section 5.4 of this SER.

4.7.4 NuScale SIET Steam Generator Tests

NuScale conducted HCSG experiments at SIET laboratories, in Piacenza, Italy. The experiments were done to evaluate the heat transfer capability of the NuScale HCSG and develop the NuScale specific model. However, the LOCA analysis does not credit heat removal from the SG. Therefore, the detailed review of the SIET test and relevant assessments is documented in the NRC staff’s SER of the Non-LOCA EM TR (ML20042E039).

4.7.5 NuScale NIST-1 Test Assessment Cases

The NuScale Power Module is dramatically different from other typical light water reactors. As discussed in Section 7.5, “NuScale NIST-1 Test Assessment Cases,” of the LOCA TR, NuScale built the NIST-1 test facility at Oregon State University to obtain test data relevant to its unique NPM design and approach to LOCA evaluations. NIST-1 was designed to model the major components of the NPM: the NPM at 1:3.3 length scale, 1:227.5 volume scale, and 1:1 time scale. NuScale performed a number of NIST-1 tests to assist in validation of the NRELAP5 system thermal-hydraulic code, and the NRC staff reviewed the summarized test information in the LOCA TR. Further, the NRC staff performed a detailed audit of the test assessments, as described in the associated audit report (ML20010D112).

Test Facility

NuScale built the NIST-1 test facility to model a scaled representation of the NPM major components with minimum distortions relative to the actual NPM in order and provide the measurements necessary for validation of the NRELAP5 model used for LOCA safety analysis. Figure 7-75, “Schematic of NuScale integral test facility and NRELAP5 nodalization,” of the LOCA TR provides a schematic of the NIST-1 facility. Even though NuScale attempted to minimize the distortions between the NIST-1 scaled test facility and the NPM, the NRC staff notes that distortions cannot be eliminated. Therefore, the NRC staff evaluates the NIST-1 facility design and tests for NRELAP5 code evaluation against important LOCA phenomena and not as testing to directly evaluate the safety or acceptability of the NPM design.

The NRC staff finds that one of the significant distortions of NIST-1 relative to the NPM is that the NIST-1 representation of the RPV, is not contained within the NIST-1 representation of the CNV. In addition, the NIST-1 representation of the CNV is not immersed in the NIST-1 41

representation of the cooling pool vessel (CPV). The NIST-1 models of the RPV and CNV are separate vessels connected by piping. Figure 7-75 shows how the NIST-1 valves that represent the RRVs and RVVs and potential pipe breaks enables the NIST-1 facility to measure flow through this piping. The NIST-1 representation of the CNV is connected to the NIST-1 representation of the CPV through a heat transfer plate (HTP). The size of the NIST-1 HTP is scaled to represent energy transfer from the entire NPM CNV inside surface to the pool.

To approximate the NPM natural circulation flow, the NIST-1 test facility represents the NPM nuclear fuel with electrically heated rods. These NIST-1 electrically heated rods establish a natural circulation flow up through the riser to the NIST-1 SG and then back to the core like the NPM design. The NIST-1 system pressure is controlled by the pressurizer component which contains heater rods to bring the pressurizer fluid up to saturation temperature at the design system pressure.

As described in Section 7.5, “NuScale NIST-1 Test Assessment Cases,” of the LOCA TR, data from the NIST-1 facility is used for both integral and separate effects validation of various phenomena. The NRC staff reviewed the NIST-1 facility design and determined that the areas of potential distortions were appropriately identified and the impact on the assessment results. One area that the NRC staff focused its review on, was the reduction of the CNV distortion. For certain NIST-1 tests, the NIST-1 CNV is preheated to reduce the distortion between the NIST-1 CNV arrangement and the actual NPM design. In NIST-1, [[

]].

As described in the associated audit report (ML19282C504), the NRC staff audited NuScale’s NRELAP5 sensitivity studies used to establish the temperature criterion for preheating. NRELAP5 computations compare the impact of [[

]]. The NRC staff concludes that CNV preheating provides an acceptable approach to minimizing the distortion introduced by the separated CNV vessel in NIST-1 for the reasons specified above. NIST-1 Integral Effects LOCA Test Procedure

As described in Section 7.5.1.6, “Steam Generators,” of the LOCA TR, a valve and switch lineup is performed to configure the NIST-1 facility for each test. The NIST-1 line modeling the LOCA break location specified for the test, is connected between the RPV and its associated CNV penetration. Orifices with the specified diameters are installed in the RVV and RRV lines to model the number of valves that are to open when ECCS actuates. The NIST-1 facility operates at a lower pressure than the NPM, and the fluid masses are scaled. The NRC staff audited the facility and test procedures, as described in the associated audit report, and conducted QA inspections. These inspections are documented in inspection reports dated October 7, 2017 (ML15268A186) and July 24, 2017 (ML17201J382).

The NRC staff understands the limitations of the NIST-1 facility, which NuScale has accounted for in its test procedures. However, the NRC staff notes that these differences between NIST-1 and the NPM mean that a limited direct comparison can be made between NIST-1 tests to NPM

LOCA results. The NRC staff notes that NIST-1 is a test facility for LOCA code development and that NIST-1 is the only facility that closely represents an NPM to simulate a LOCA.

NIST Facility NRELAP5 Model

The NRC staff reviewed and audited details about NuScale’s NRELAP5 nodalization model, as described in the associated audit report (ML20010D112), which is similar to the model used for the NPM and is described in Section 7.5.2, “Facility NRELAP5 Model,” of the TR. The NRELAP5 model is a complete one-dimensional representation of the NIST-1 test facility. The NRC staff finds that the NuScale NIST-1 NRELAP5 model provides an acceptable representation of the NIST-1 test facility to evaluate the capability of NRELAP5 to model NIST-1 tests.

NIST Facility Test Matrix

The NRC staff reviewed NuScale’s test matrix, given in Table 7-6, “Facility high priority tests for NRELAP5 code validation,” of the LOCA TR and finds the suite of tests are sufficient to benchmark the NRELAP5 computer code and justify its use for LOCA analyses. Each of this series of tests is evaluated below regarding its applicability to NuScale.

Separate Effect High Pressure Condensation Tests

NuScale performed Test HP-02 to assess the capability of NRELAP5 to predict condensation rates at high pressure test conditions. While HP-02 was a quasi-steady test, a transient was performed to achieve the desired steady-state test conditions. The HP-02 test included direct measurements of the CNV pressure, CNV level, CNV temperature, and CPV temperature response. The CNV was a closed vessel, so, condensed steam (water) accumulated to produce a rising liquid level. Details of the test procedures are presented in the LTR, Section 7.5.4, “Separate Effect High Pressure Condensation Tests.”

NuScale reports generally good agreement between NRELAP5 and test data for CLL, upper containment wall temperature and upper containment wall temperature at the cooling pool. The reported pressure is over-predicted by NRELAP5 as the absolute pressure increases. This is an important test because it is the only “larger” scale test available to validate the Extended Shah condensation modeling in NRELAP5. As described in the associated audit report (ML19282C504), the NRC staff audited NuScale’s test assessments and reviewed information regarding the pressure overprediction (ML18256A361) and noted that the primary cause of over-predicting the HP-02 peak pressure conditions is due to how NRELAP5 calculates the condensate film thickness when two heat structures connected to a single volume are both acting as condensing surfaces. The HP-02 prediction of peak pressure is affected by this code limitation during the initial containment pressurization because both the shell wall heat structure and HTP heat structure are initially cold.

NuScale performed sensitivity calculations and the user input heated hydraulic diameter was modified based on the CNV shell and HTP geometry to account for the code treatment of liquid in a condensing volume when calculating film thickness. The sensitivity calculations result in a significantly improved prediction of the pressure rate of increase and peak pressure. The results of these sensitivity calculations show that within the pseudo-steady state period, NRELAP5 can predict well the condensation and heat transfer rates when a single surface is the dominant condensing surface. The limitation arises when there is more than one surface 43

with significant condensation connected to a hydraulic cell. Once the CNV shell has ceased participation in the condensation process, the assumption that only one dominating condensing surface exists becomes valid.

The NRC staff’s review recognizes that the two condensing surfaces can distort the condensation processes for HP-02; however, the adjustments made to code input only applies to the initial heatup. Therefore, the use of the laminar film condensation correlation would result in an under-predicted film condensation coefficient and over-predicted pressure. The NRC staff has determined that the current NRELAP5 computation of pressure in HP-02 demonstrates a conservative computation of pressure.

The NRC staff also reviewed data and NRELAP5 overlay boundary condition plots for inlet steam flow, pressure, and inlet steam temperature (superheated), as well as plots for CLL and condensation rate (ML18256A361). The NRC staff notes that while there is a general agreement between the temperature data and NRELAP5 computations, the CNV fluid temperatures at the lower elevation are noticeably over-predicted.

NuScale noted that the overall comparisons between the NRELAP5 results and data indicate a reasonable to excellent agreement. Considering instrumentation and other measurement uncertainties, the NRC staff considers the results to be reasonable. The cause of the containment pressure over-prediction in the HP-02 test is the NRELAP5 treatment of film thickness when two heat structures connected to a single volume provide condensation surfaces.

The NRC staff noticed that most of the Reynolds numbers are reported near 1000. The laminar film regime ends at a Reynolds number of about 30 and enters the wavy-laminar regime out to a Reynolds number of about 1800, and that regime has higher condensation heat transfer coefficients. Thus, under-predicted film condensation coefficients could contribute to the over- predicted pressure in HP-02 Run 3.

The applicant reported (ML18256A361) that the HTP thermal conductivity was estimated from [[

.]]

The NRC staff reviewed the results of NRELAP5 using radial nodes of 23, 29 (base), and 36, provided by the applicant (ML18256A361). The results demonstrate that there is no dependence between the CNV pressure response and the three nodalization schemes investigated. The temperature profile through the HTP from the CNV to CPV shows minimal changes between the sensitivities, and, there is no discernible difference in the integrated condensation rates.

The NRC staff reviewed the capability of the NRELAP5 code to adequately represent thermal stratification in the NIST-1 CNV and the applicant’s justification that the validation of condensation is accurate. The NRC staff reviewed the impact of node size on interfacial condensation at the steam-water interface in HP-02 and agrees with the applicant that node size had a small impact on the computed interfacial condensation and pressure.

As described in the associated audit report (ML19282C504), the NRC staff audited NuScale’s assessment calculations that applied an adiabatic boundary condition to assess the impact of including shell wall heat losses on the containment pressure response in HP-02. The NRC staff noted that the applicant’s sensitivity calculation results show [[ ]].” The NRC staff also audited (ML19282C504) the NIST-1 HP-49 test nodalization sensitivity analyses of lower CNV 49 noding for NIST-1 test HP-49 that considered an inadvertent RRV opening. Three node size configurations were considered – coarse, base and fine. The NRC staff confirmed that the NRELAP5 model sensitivity calculation results for a NIST-1 inadvertent RRV transient event show that further refined or coarser CNV nodalization has an insignificant effect on the containment pressurization response compared to the base nodalization. Natural Circulation Test at Power

NuScale performed NIST-1 test HP-05, to assess the capability of NRELAP5 to predict natural circulation flow at various core powers and test conditions by comparing experimental data and NRELAP5 predictions. The specifics of the test configuration are described by NuScale in its LOCA TR and the NRC staff reviewed these descriptions.

NuScale calculated form losses for a base run using Idelchik as referenced in the LOCA TR for the various geometric configurations around the loop. NuScale modified the losses based on the individual experimental differential pressures measured around the flow loop and then confirmed the global response by comparing the experimental loop flow rate to that predicted by NRELAP5.

The NRC staff reviewed the comparisons illustrated in the LOCA TR and agrees that NRELAP5 is capable of predicting primary flow rate, core inlet temperature, and core outlet temperature with a reasonable-to-excellent agreement for natural circulation flow conditions. The NRC staff audited the supporting testing reports and the measured data from Test HP-05, as described in the associated audit report (ML20010D112), and observed that it was a calibration test to refine hydraulic loss coefficients to improve the NRELAP5 computation of natural circulation flow rates observed in NIST-1. The NRC staff reviewed the LOCA base deck and noted that the RCS form losses, were only based on theoretical formulations from Idelchik for flow regions outside of the core and SG.

NuScale tests showed that major pressure losses are in the reactor fuel and SG and that they are well characterized by fuel vendor testing and large scale HCSG testing at SIET Laboratories, in Piacenza, Italy. As the formulas from Idelchik are widely accepted for single phase flow and the dominant loss through the NPM RCS is from the core and SG, which were well characterized by fuel bundle and SG head loss testing, the NRC staff found that the HP-05 test supported the conclusion that the NRELAP5 code is capable of predicting the natural circulation of an NPM.

45 Chemical and Volume Control System Loss-of-Coolant Accident Integral Effects Tests

NuScale performed Test HP-06 and HP-06b, to assess the capability of NRELAP5 to predict the integral response and multiple phenomena of the NIST-1 facility for a single-ended discharge line break inside containment. The specific test conditions and configuration are detailed in Section 7.6.5 of the LOCA TR, which the NRC staff reviewed. NuScale compared several parameters to assess an agreement with NRELAP5, including: direct measurements of the CNV pressure, RPV pressure, CNV level, RPV level, primary flowrate, break orifice differential pressure, pressurizer level, CPV temperature, CNV temperature, and HTP temperature.

For both tests, NuScale reports that the comparison between the calculated and measured results are in a reasonable-to-excellent agreement. The NRC staff audited the supporting test reports, as well as an evaluation of the NIST-1 HP-06 test results and the impact of preheating of the NIST-1 containment, as described in the associated audit report (ML20010D112), which provided a justification to show how the NRELAP5 code correctly calculates the temperature, enthalpy and mass fraction of vapor and liquid as the containment pressure increases with time.

The experimental data plot that the NRC staff audited, as documented in the associated audit report (ML20010D112), showed the measured liquid temperatures versus time at four elevations. The plot clearly shows the accumulation of thermally stratified subcooled water in the presence of walls that have been preheated. The temperature response shown by the CNV thermocouples matches the expected thermal stratification trends. No adverse effects due to preheating are observed. Thus, the NRC staff agrees with the applicant that the test is judged to be adequate to assess the ability of NRELAP5 to model the thermal stratification phenomenon.

As discussed in the associated audit report (ML20010D112), the NRC staff also audited an analysis of condensation at the steam-water interface in the CNV (pool condensation). Physically, the thermal stratification of the CNV pool has the effect of limiting surface condensation, particularly during the phase when the containment pressure is increasing. The analysis indicates that an upper bound on the pressure error due to over-prediction of pool condensation is less than one percent for the HP-06b test. The applicant stated that given the small impact of pool condensation on pressure results, the nodalization is appropriate for purposes of modeling pool condensation. The NRC staff found that this information provides sufficient justification for applying the NRELAP5 pool condensation model.

Assessment of NRELAP5 Prediction of Peak Containment Pressure

The NRC staff notes that NuScale relied on the NRELAP5 LOCA methods to perform peak containment pressure analysis. Figure 7-102, “NIST-1 HP-06b containment vessel pressure comparison,” of the LOCA TR, showed that NRELAP5 slightly overestimated the measured NIST peak containment pressure with a negligible deviation. As discussed in the associated audit report (ML20034D464), the NRC staff audited information that indicated that there were uncertainties in the NIST containment pressure measurement instrumentation and core heater rod center line thermocouple readings. In addition, NuScale identified the uncertainties associated with the NRELAP5 NIST-1 model initial and boundary conditions. The applicant used this revised NIST-1 modeling in its assessment report (ML18268A365) regarding HP-49 RRV opening test results.

46 NuScale evaluated the heater rod model uncertainties using three recently completed tests, which included the HP-43, the NLT-15p2, and the HP-49 tests. The NIST-1 RPV core is made up of electrically heated rods, some of which are fixed with internal thermocouples. Each heater rod contains a heater element that is inserted into a thermowell where a nominal 0.005-inch gap, between the heater element and the thermowell, is completely filled with boron nitride to maintain sufficient heat transfer within the heater element to moderate heater element temperatures. The heater rods then are seal welded at the top but remained open at the bottom. NuScale explained that over time, the gap of some of those heater rods may have lost some of the boron nitride resulting in higher element temperatures and higher initial stored energy than when they were newly installed. The applicant’s initial base NIST-1 NRELAP5 model included a uniformly applied rod model with no fixed air gap in the rods for all NIST-1 tests. This approach resulted in a potential underestimation of initial rod stored energy, and according to NuScale, accounted for under prediction of CNV pressure in the early test assessments. In its examination of the HP-43 and HP-49 test data, during the Containment audit, as discussed in the associated audit report (ML19282C504), and QA inspection (ML19093A669), the NRC staff found that NuScale used the maximum of measured temperature data rather than averaged values. Although NRELAP5 inputs should be based on average temperature where data is available and that conservative input should be used for older previous tests where rod temperature data was not collected, the NRC staff also noted that it is likely that the boron nitride layer eroded with time as more tests were completed, suggesting that lower heater element temperatures would be more realistic, especially for the earliest tests.

The NRC staff performed sensitivity studies and determined that differences in the results with the lower realistic initial temperature were minimal and not large enough to affect the overall conclusions of the assessment. Therefore, the NRC staff found the applicant’s analysis and rod modeling to be acceptable. The results show a conservative over-prediction of containment pressure by approximately 10 to 12 psi, therefore, there is sufficient margin to conclude that NRELAP5 adequately predicts peak CNV pressure for NIST-1 facility tests. Therefore, based on the review of HP-49 test results and the re-analyses of HP-06, HP-06b, HP-07, HP-09 and HP-43 using the revised NRELAP5 NIST-1 model, the NRC staff found the assessment results to be acceptable to justify the use of the NRELAP5 code to perform peak containment pressure analysis.

Pressurizer Spray Supply Line Loss-of-Coolant Accident Integral Effects Test

NuScale performed the HP-07 test benchmark to assess the capability of NRELAP5 to predict the integral response of the NIST-1 facility modeling a single-ended pressurizer spray supply line break inside containment. The phenomena evaluated in the HP-07 test were the same as those in the HP-06 test. The NRC staff reviewed the applicant’s comparisons, which showed the comparison between that the calculated and measured results are in a reasonable-to- excellent agreement. The NRC staff agrees that they are in a good agreement and the assessment is therefore, acceptable.

Spurious Reactor Vent Valve Opening Test

NuScale performed the HP-09 test to assess the capability of NRELAP5 to predict the integral response of the NIST-1 facility modeling to inadvertent depressurization of the RPV initiated by a spurious opening of an RVV without DHRS. Furthermore, this test also provided a bounding depressurization rate for a LOCA initiated by break from pressurizer gas space. 47 The phenomena evaluated in the HP-09 test were the same as those in the HP-06 test. The NRC staff reviewed the comparisons provided in the LOCA TR, which include core power, RVV mass flow rate, RCS pressure and level, containment pressure and level. The NRC staff agrees that the comparison between the calculated and measured results are in a reasonable-to- excellent agreement and the assessment is therefore, acceptable.

4.8 Assessment of Evaluation Model Adequacy

In Section 8, “Assessment of Evaluation Model Adequacy,” of the LOCA TR, NuScale presented its assessment of the adequacy of its LOCA EM based on the NRELAP5 computer code Version 1.4 and Revision 2 of the NPM plant base model for analysis of design-basis LOCAs. NuScale demonstrated LOCA EM adequacy by closure model and correlation reviews, and assessments against relevant experimental data. The NRC staff focused its review on being consistent with the EMDAP (RG 1.203).

4.8.1 Adequacy Demonstration Overview

Section 8.1, “Adequacy Demonstration Overview,” of the LOCA TR provides a summary of the NuScale process for demonstrating model adequacy. NuScale used the results of its PIRT process discussed in Section 4 of the LOCA TR, to select the important phenomena for demonstrating LOCA model adequacy. The NRC staff’s findings on the NuScale LOCA EM are provided below for each of NuScale’s adequacy determinations.

4.8.2 Evaluation of Models and Correlations (Bottom-Up Assessment)

As discussed in Section 8.2, “Evaluation of Models and Correlations (Bottom-Up Assessment),” of the LOCA TR, NuScale evaluated the adequacy of NRELAP5 for modeling the PIRT high ranked phenomena by comparing NRELAP5 analyses against appropriate fundamental and special effects data. As discussed further below, the NRC staff reviewed NuScale’s process for selecting fundamental and special effects test data to evaluate its LOCA EM for highly ranked PIRT phenomena and finds it to be acceptable because it conforms to the process described in RG 1.203.

Important Models and Correlations

The NRC staff reviewed NuScale’s identified high ranked PIRT phenomena and the dominant NRELAP5 models and correlations required to assess these phenomena, as well as the key parameters, special situations associated with the phenomena and NRELAP5 assessments with NuScale and legacy test data used. As part of its review, the NRC staff reviewed information (ML17310B505) provided by NuScale explaining the validation of the modeling of flow through the ECCS valves. The NRC staff observed that [[

48 ]]

The NRC staff assessed the adequacy of NRELAP5 for modeling interruption of natural circulation. The NRC staff reviewed information provided by NuScale (ML18031B319) to demonstrate the adequacy of modeling the core as two parallel channels without crossflow, including the applicant’s computational assessment of crossflow modeling and CHF computations using VIPRE-01. The NRC staff agrees with the applicant’s conclusions that the VIPRE-01 computations show that allowing full crossflow (base case), produces a higher flow rate and an associated lower void fraction in the hot assembly than for the other two restrictive crossflow models. The result is a larger CHFR when crossflow is allowed. The reported results support the applicant’s conclusion that the closed channel model in NRELAP5 produces a more conservative CHF margin than an open channel model. Therefore, based on the above discussion, the NRC staff concludes that the modeling in NRELAP5 for the interruption of natural circulation is sufficient.

The NRC staff reviewed the applicant’s estimated range of key NPM steady-state and design basis LOCA parameters that NuScale used to evaluate the adequacy of its LOCA EM TR models and correlations. NuScale stated that these parameter ranges shown in Table 8.2, “NuScale Power Module range of process parameters,” of the LOCA TR identify the minimum range for demonstrating NRELAP5 adequacy, but that the applicability of models and correlations are not restricted to these ranges. NuScale determined that these parameter ranges from several sources including design values, proposed technical specification limits, and limiting initial and boundary conditions. NuScale obtained the ranges for some parameters from the NRELAP5 LOCA break spectrum calculations described in Section 9.0, “Loss-of- Coolant Accident Calculations,” of the LOCA TR. The NRC staff finds that the NuScale process used for determining parameter ranges is acceptable because the values are based on the design, or are conservative, or are limited by technical specifications.

Two-Phase and Single-Phase Choked Flow (Mass and Energy Release)

As discussed in Section 4.6.6.1 of this SER, NRELAP5 employs a critical flow model that uses the [[

]]

The NRC staff reviewed NuScale’s comparison of this model to [[

49 .]]

The NRC staff agrees with NuScale’s conclusions that the NRELAP5 comparisons to these two tests demonstrated a good agreement with the data during the subcooled portion of the tests, while over-predicting the break flow for saturated conditions, thereby displaying a conservative prediction of break mass flow rate. The NRC staff finds that these critical flow tests comparisons are sufficient to demonstrate an acceptable performance supported by the finding that the [[

]]

[[ ]]

The NRC staff reviewed the applicant’s comparison of the NRELAP5 model for [[

]] The NRELAP5 code predicted-versus measured-pressure drop, of Figure 7-2, “Predicted versus measured pressure drop for selected contraction tests,” of the LOCA TR, [[

]] showed an overall acceptable agreement. [[

]] The NRC staff reviewed the comparison results and agrees that the values are conservative. Therefore, the NRC staff finds this modeling approach to be acceptable.

[[ ]]

[[ ]] The NRC staff reviewed Table 8-6, “Dimensions of NuScale Power Module, NIST-1 and Bankoff pressurizer plate,” in the LOCA TR, which indicates that [[

.]]

The NRC staff noted that the verification of the correlation is based on the fact that the pressurizer drainage is well predicted in the test simulation. Moreover, the NRC LOCA verification runs with and without CCFL indicate that the results are not sensitive to CCFL at the pressurizer baffle plate or the core upper plate. Furthermore, regarding the limiting small break LOCA, all potential CCFL effects will have subsided due to the very low steaming rate at the time that the minimum liquid level is reached late in the event, where the liquid level and hence two-phase swelled level, in the vessel remains well above the top elevation of the core. [[

]] Review of a CCFL paper by Stephen and Mayinger, “Experimental and Analytical study of Countercurrent Flow Limitations in Vertical Gas Liquid Flows,” Chem. Eng. Tech. 15 (1992) pp 51-62, shows comparisons of the Wallis and Kutateladze forms of which Bankoff is intermediate, to a range of pressure flooding conditions up to about 200 psia. They importantly noted that the reducing effect of high gas-phase densities on gas velocities during flooding was satisfactorily predicted by these correlations.

Based on the NRC staff’s assessment, as discussed above, the NRC staff finds [[ ]] to be acceptable. [[

]]

Flashing

The NRC staff imposes no special findings or limitations on the modeling of NPM due to the flashing models implemented in the NRELAP5 code. NRELAP5 is judged by the NRC staff to 52 be able to predict flashing during a depressurization event. This is irrespective of the nonconservative behavior shown by the interfacial drag model to accurately predict void and level swell separate effects test, where the model tends to over-predict two-phase levels and void distribution behavior in the axial direction.

[[

.]] [[

]] [[

.]] [[

]] 4.8.3 Evaluation of Integral Performance (Top-Down Assessment)

The NRC staff reviewed NuScale’s primary areas described in LOCA TR Section 8.3, “Evaluation of Integral Performance (Top-Down Assessment),” of its top-down assessment, including the code governing equations, numerical solution and underlying assumptions; the integrated performance of the code against the IETs conducted in the NIST-1 facility; and calculations to evaluate differences and distortions between the NIST-1 facility and the NPM design.

The NRC staff reviewed the NRELAPS NIST-1 and NPM input models and found that these models were developed using generally consistent nodalization and option selection and that the IETs at the NIST-1 facility are appropriate for evaluating the key phenomena for NPM LOCA analyses.

Review of Code Governing Equations and Numerics

The NRELAP5 code governing equations and numerics are described in the LOCA TR, Section 8.3.1, “Review of Code Governing Equations and Numerics,” and are the same as that of the original RELAP5-3D code. Therefore, the NRC staff determined a further in-depth review was not necessary. The NRC staff’s review of the hydrodynamic model and field equations were discussed earlier in Sections 4.6.2 of this SER.

NuScale Facility Scaling

The NIST-1 facility is designed to simulate the integral system behaviors of a single NPM immersed in a reactor building pool. The NRC staff audited the applicant’s scaling analysis, as discussed in the associated audit report (ML20034D464), to determine the NIST-1 dimensions and operating conditions. Distortion between the test facility and prototype is also analyzed in the scaling analysis. The hierarchical two-tiered scaling (H2TS) methodology was adopted by NuScale to scale the phenomena including RCS natural circulation, LOCA progression and ECCS operations. The NRC staff focused its review on confirming that non-dimensional numbers (π groups) representing phenomena are preserved for both the NPM and NIST-1 to capture the high-ranked phenomena identified in the LOCA PIRT. The FOMs in the NuScale design are the minimum CLL in the core, the peak CNV pressure and CHF.

NIST-1 has inherent distortions due to its small size and different component layout compared to NPM. Distortion also arises due to the difference of operating conditions in specific transients. The scaling analysis covered CVCS LOCA transient (HP-06) and additional LOCA scenarios: the high-point vent line break (HP-07) and inadvertent opening of RVV (HP-09). These break locations cover reactor coolant liquid space and vapor space. The NRC staff audited the applicant’s NRELAP5 analyses for NIST-1 and the NPM, including its evaluation of scaling distortions and their impacts on FOMs. 57

Hierarchical Two-Tiered Scaling is a proven methodology developed by the NRC and has been used in several reactor designs. The NRC staff reviewed the scaling summary in the LOCA TR, supplementary information provided by the applicant (ML19058A867) and audited the details of the implementation of this methodology in the NuScale scaling analysis to determine whether it is appropriately used.

The NRC staff audited both stages of the applicant’s scaling analyses. The first stage was steady state single-phase natural circulation in the RPV. In NIST-1, the maximum power level is scaled [[ ]] The NIST-1 vessel dimensions were determined in this stage. In the second stage, the applicant performed scaling on LOCA phenomena at different phases. Potential distortions were analyzed and identified through the difference in non-dimensional π groups.

The NRC staff reviewed the four groups of transient phenomena NuScale analyzed, including: vessel depressurization and containment pressurization during the blowdown and venting phases, the long-term recirculation phase and reactor building pool heat up. The NRC staff found that NuScale’s scaling analyses correctly identified the control volumes of interest, the interactions between components and phases of event progression.

STEADY STATE NATURAL CIRCULATION OPERATION SCALING

The NRC staff reviewed the scaling ratios of the NIST-1 and NPM dimensions. The facility was designed to preserve event time and power-to-volume ratio. The NRC staff’s audit focused on four areas of interest: the downcomer to lower plenum flow path, the central core region, the flow path between the upper riser and annulus, and the SG external flow. Among these, the SG frictional pressure losses dominate. The NRC staff found that NuScale applied appropriate scaling factors and initial steady state conditions with buoyancy forces balancing frictional losses, resulting in the correct flow rate comparisons. The one-to-one time ratio (isochronicity) requirement was met. Based on the analysis, the NIST-1 facility was designed to have a much higher loop resistance than the NPM. The NIST-1 SG scaling and the derivations of non- dimensional π groups for steady state natural circulation were confirmed by NuScale using additional NRELAP5 analysis with an excellent agreement of flow predictions and data.

The NRC staff reviewed single-phase natural circulation analyses for the NPM at different powers (100 percent and 50 percent) and for NIST-1 at different pressures and confirmed that there is not much effect from the pressure on the NIST-1 natural circulation flow. Therefore, the distortion due to lower pressure in the NIST-1, does not impact the results of natural circulation scaling. The matching of the natural circulation number and loop energy ratio lead to correctly scaled flow rates in NIST-1 compared to the NPM at 50 percent rated power. In the 100 percent NPM power condition, the scaled flow ratio and core temperature’s distribution are slightly different than those in the 50 percent power condition. However, as the LOCA starts, the phenomena are the same if the decay power is scaled from 100 percent NPM decay power.

The NRC staff also performed TRACE and NRELAP5 confirmatory calculations and confirmed that the appropriate NIST-1 loop resistance was established to confirm the one-to-one time ratio requirement. Therefore, the NRC staff concludes that the scaling analyses for natural circulation are acceptable.

LOSS OF COOLANT ACCIDENT AND EMERGENCY CORE COOLING SYSTEM SCALING

Section 8.3.2.3, “Loss-of-Coolant Accident and Emergency Core Cooling System Scaling,” of the LOCA TR, summarizes the scaling of vessel depressurization, containment pressurization, long-term recirculation and building pool heat up phenomena for the CVCS line break event. The NRC staff audited the scaling approach detailed in NuScale’s scaling reports, as discussed in the associated audit report (ML20034D464), and concluded that it is correct in terms of identifying control volumes and phases of LOCA progression.

Vessel Depressurization NuScale’s scaling formulation includes vessel mass and energy balance equations. [[

.]] Therefore, the NRC staff determined that NuScale’s scaling analyses for this phase is acceptable.

Containment pressurization [[

]]

Long Term Recirculation Cooling Phase The NRC staff reviewed NuScale’s top-down and bottom-up scaling flow path and notes that it correctly identified important phenomena that controls the steam flow and the return of condensate. The NRC staff also reviewed NuScale’s containment pressure equation, which is 59

formulated with flow out of the vessel through the break and RVV, flow back into the vessel at the RRV, and heat loss to the pool. During this phase, the pressure drop between the RPV and CNV is determined by flow resistance and the flow rate. As the actual NIST-1 hydro-static driving head was scaled less than that of the NPM, the flow resistance in NIST-1 was evaluated to confirm the resistance ratio. For the CNV inventory balance, the two important phenomena are RRV flow and CNV wall phase change. RRV flow shows a large distortion, but in the acceptable range. For the energy balance, the most important phenomenon is the CNV wall heat transfer and the next important phenomenon is RVV energy flow. The distortions in both phenomena are less than 15 percent and are considered to be insignificant by the NRC staff. The NRC staff audited NuScale’s calculations, as discussed in the associated audit reports (ML20034D464), and concluded that the scaling formulations and approaches for this phase are appropriate because there is less than 15 percent distortion.

Building Pool Heat up The NRC staff audited NuScale’s scaling related to the ultimate heat sink. NIST-1 has a HTP that connects the CNV with the pool, and the pool is a separate tank with its volume scaled as the power of the reactor for only one bay of the common pool. Therefore, the natural circulation pattern in NIST-1 is different than the multi-module pool in the NPM. Less horizontal thermal diffusion is expected. NuScale recognized the complexity of mixing behaviors in the stratified layer near the CNV wall but did not include the scaling of diffusion flows. The approach of neglecting the thermal diffusion in scaling is conservative since the diffusion helps cooling, and the diffusion flow will eventually reduce as the pool warms up after the initial period. Because the approach is conservative, the NRC staff finds it acceptable.

HIGH VENT LINE BREAK (HP07) AND INADVERTENT OPENING OF ONE RVV (HP09)

In addition to the CVCS break LOCA, the applicant performed scaling analyses for two other events: High Vent Line break (HP07) and inadvertent opening of one RVV (HP09).

The NRC staff reviewed the values of scaling groups at four snapshots in time during the high point vent line break (HP-07). The snapshots are: 1) when RPV pressure reaches 1500 psia, 2) near peak pressure, 3) right before ECCS actuation, and 4) long term cooling with RRV flow reversal. The NRC staff noted that the inventory is controlled by flashing and there is negligible distortion in flashing phenomenon for the RPV. The energy balance involves three important phenomena, break energy flow, core heat transfer and SG heat loss. The largest distortion is in SG heat transfer, but it is considered by NuScale to be small compared to break energy. In CNV scaling, the NRC staff noted that the scaling groups are small and the transient is mild, as compared to HP-06. The NRC staff concluded that in general, the dominant phenomena are well represented by NIST-1, as distortions are small in this transient.

HP09 is a fast transient, and the NRC staff reviewed the applicant’s estimates for three snapshots in time: initial pressure, peak CNV pressure and long-term cooling.

[[

.]]

The NRC staff notes that even acceptable scaling distortions can have a cumulative effect on FOMs. The NRC staff observed this during its audit of NuScale’s distortion report, as described in the associated audit report (ML20034D464). However, as the scaling groups for important phenomena were in the same order of magnitude for both NIST and the NPM, the NRC staff finds that the data are appropriate for code validation. The impact of scaling distortions is discussed below.

Assessment of NuScale Facility Integral Effect Test Data

The NRC staff reviewed NuScale’s summary of NIST-1 IET tests that support NPM calculations and its assessment of data in these tests: HP-05, HP-06, HP-07, HP-09, HP-43 and HP-49, with NRELAP5. NuScale performed a sensitivity study for these tests to evaluate the potential uncertainties. NuScale used the agreement between the prediction and test data for these tests to demonstrate the applicability of NRELAP5 in modeling high-ranked phenomena in the NuScale design.

Evaluation of NuScale Integral Effects Tests Distortions and NRELAP5 Scalability

The NRC staff reviewed the applicant’s summary of distortions concerning biases of initial/boundary conditions (IC/BC) and audited its actual operating procedure for the as-built 61

NIST-1 facility. The NRC staff audited the applicant’s NRELAP predictions for tests HP-05, HP-06, HP-07 and HP-09. In addition, the NRC staff audited the applicant’s three NPM sensitivity calculations for each transient (base case for proposed design without Appendix K assumptions, IC/BC case with the same initial and boundary conditions as in NIST-1, and a distortion case including scaling distortions). The NPM results were adjusted using scaling factors before comparing to NIST-1 data. The NRC staff audited the applicant’s NRELAP5 prediction of NIST-1 tests with NIST-1 data, as described in the associated audit report (ML20010D112). The NRC staff concludes that the summary in the LOCA TR accurately described the distortions, which did not invalidate the scaling analysis results.

[[

]]

The NRC staff concludes that, because of the size of the NIST-1 facility and its development history, there are scaling distortions. These distortions have been identified by the applicant and its impact on the FOM have been fully assessed and understood. The test results from NIST-1 were not only used to assess NRELAP-5 code, but also can be representative for NPM during LOCA and IORVs.

4.8.4 Summary of Adequacy Findings

The NRC staff reviewed the adequacy of the NRELAP5 code for analysis of design-basis LOCAs in the NPM and focused on the NuScale identification of key phenomenological models in a PIRT that are needed to successfully predict ECCS performance following a LOCA. This is demonstrated by choosing the proper closure models and correlations, and then assigning the many assessments against relevant separate effects tests and integral experiments to validate the important models listed in the PIRT. The NRC staff considers this a key step in establishing the adequacy of the NRELAP5 code as an acceptable component of the NuScale LOCA methodology as part of the EMDAP given in RG 1.203. The NRC staff also reviewed the subsequent steps to this objective, including documentation of the bottom-up assessment of the NRELAP5 models and correlations to determine their adequacy to predict the high (H) ranked phenomena in the PIRT, as well as a top-down assessment of the EM, including a review of EM governing equations and numerics to determine their applicability to NPM LOCA analysis, and evaluation of the integral code performance based on the assessments of the EM against relevant IETs. The NRC staff reviewed the applicant’s summary in the LOCA TR of the adequacy findings, which showed how each PIRT high (H) ranked phenomenon is covered by the LOCA methodology models and correlations. The NRC staff also reviewed the applicant’s identification of models that are marginally adequate, or ranges in which PIRT phenomena are not covered, and the manner of compensating for code limitations.

The LOCA TR identifies key models and correlations which are important to predicting the NPM LOCA ECCS performance following a LOCA. These are listed in Section 8.2, “Evaluation of Models and Correlations (Bottom-Up Assessment),” of the LOCA TR. The NRC staff noted that the list does not contain Baker Just for oxidation nor a rod swelling and rupture model since the acceptance criteria for acceptable ECCS performance does not include core uncovery. The Henry/Fauske Moody critical flow model meets Appendix K requirements, as well as the decay heat model, which uses the 1973 ANS standard with the 1.2 multiplier and inclusion of actinide decay.

The NRC staff notes that for current generation plants, downcomer boiling is an important phenomenon affecting the event progression following large break LOCAs. Because large breaks are not possible for NPM due to the design and the quick cool down of reactor vessel wall, the NRC staff agrees with NuScale that the downcomer boiling in NPM is not significantly large enough to produce a lower long-term liquid level in the core and riser region.

The NRC staff also agrees that these are relevant phenomenological models for simulating small break LOCAs in the NuScale NPM. Further, the NRC staff believes that the code predictions of the basic phenomena, such as these, with behavior observed in single situations created in individual separate effects test facilities allow a more focused and better assessment of the accuracy of the specific models in the code to be made than is possible using integral experimental data. This is because separate effects tests are dedicated to the study of a single particular phenomenological characteristic, so the measurement instrumentation can also be chosen more appropriately. 64

The 21 dominant NRELAP5 models and correlations for LOCA modeling, listed by the applicant in Section 8.2.1, “Important Models and Correlations,” of the LOCA TR, were evaluated by the NRC staff in Sections 4.6.2 through 4.6.11 of this SER, and again summarized in Sections 4.8.2.1 through 4.8.2.22 of this SER. In brief, the NRC staff noted that the interfacial drag model was not considered accurate enough for determining the potential for core uncovery since the model over predicted the level swell and the axial void profile in many of the separate effect tests. However, the NRC staff concluded that these deficiencies would not have a significant effect on FOMs. This was noted in Section 4.6.8 of this SER as this modeling was found to be reasonably conservative in nature for the phenomenon of interest. The NRC staff concluded that the CNV condensation modeling is adequate to determine that the worst small break LOCA has been identified which displays the minimum liquid level in the core. Further, the minimum level worst case can be demonstrated to determine the liquid level above the top of the core, and the NRC staff believes the methodology is sufficient to predict the potential for two-phase uncovery of the core.

The NRC staff recognizes that there is a deficiency of integral test data against which the NRELAP5 code was validated against. And, further, there is only the NIST-1 facility that applies directly to the NPM design, which NuScale successfully compared and benchmarked the NRELAP5 code to. It is evident that the NRELAP5 modeling is capable of reproducing the NIST-1 LOCA results. From this, it is the NRC staff’s judgement that it is not unreasonable to expect that NRELAP5 is capable of producing the NPM LOCA results.

In addition, NuScale successfully applied the NRELAP5 code to two Semiscale natural circulation tests. It is also noted that this facility is a much smaller scale, and there were no specific requirements for nodalization to successfully model natural circulation. As such, given that the condensation modeling was determined to use a conservative approach, the NRC staff found that the SET and IET code qualification effort supports the acceptance of the NRELAP5 code for evaluating ECCS performance following a small break LOCA in the NuScale NPM.

The NRC staff noted that distortions may compensate and result in seemingly conservative predictions for the tests. To determine the conservativeness of an EM, the applicability and scalability of the code need to be evaluated. An assessment of the applicability of NRELAP5 based on model correlations and bottom-up phenomenon is conducted and summarized in the LOCA TR, in Table 8-18, “Summary of bottom-up evaluation of NRELAP5 models and correlations,” and Table 8-19, “Applicability summary for high-ranked phenomena,” lists 21 high ranking phenomena. The applicant justified the applicability of the NRELAP5 code based on the agreement of NIST-1 data assessment. Based on the staff’s evaluation of the applicant’s applicability assessment the NRC staff considers this approach to be acceptable.

Due to the scale-dependent correlations used in the code, a scalability evaluation of NRELAP5 models was conducted. The applicant assessed the scalability issue in Table 8-18 of the LOCA TR by examining the scale dependency of important phenomena. These scale dependent models include choked flow, CCFL model, wall film condensation, riser flow regime and 3-D core flow distribution. The applicant either performed sensitivity studies on the coefficients of the correlations (e.g. CCFL model) or used conservative assumption for scale-dependent phenomenon (e.g. laminar regime film condensation heat transfer coefficient for turbulent regime) to ensure that the FOMs are not compromised. Based on the staff’s evaluation of the applicant’s scalability evaluation, the NRC staff found the approach acceptable.

4.9 Loss-of-Coolant Accident Calculations

NuScale stated that “the primary purpose of the break spectrum calculations and sensitivity studies is to support the development of the LOCA EM and to demonstrate its application for the evaluation of the NPM ECCS performance during postulated LOCAs.”

The initial/boundary conditions and inputs for key LOCA EM parameters used for this analysis, are summarized in Appendix A of the LOCA TR.

4.9.1 Loss-of-Coolant Accident Progression in the NuScale Power Module

The NRC staff reviewed the applicant’s LOCA analyses summarized in Section 9.1, “Loss-of- Coolant Accident Progression in the NuScale Power Module,” of the LOCA TR, and audited the underlying calculations, as described in the associated audit report (ML20010D112). These calculations are for a representative liquid space break (100 percent break of the RCS injection line) and a representative steam space break (100 percent break of the high point vent line).

NuScale described representative LOCA scenarios that assume full-break area, no loss of ac or dc power, no single failure, and do not credit either decay heat removal train. The NRC staff agrees with NuScale that the representative LOCAs are appropriate to show a typical application of the NuScale 10 CFR Part 50, Appendix K LOCA EM. The NRC staff also noted that these may not be the limiting LOCA cases regarding the LOCA evaluations in Chapter 15 of the NuScale DCA, or any other application of this methodology to a specific design. Specific analytic results for LOCAs are evaluated as part of a design specific application of this methodology, such as the NuScale DCA.

Liquid Space Break

NRELAP5 calculates immediate choked flow for a 100 percent break of the RCS injection line. The mass and energy releases into the CNV through the break results in rapid pressurization of the CNV and depressurization of the RPV. The applicant described the sequence of events in Section 9.1.1, “Liquid Space Break,” of the LOCA TR. The NRC staff reviewed the applicant’s NRELAP5 calculated RPV and CNV pressure responses and noted that the NRELAP5 calculated energy release to the CNV through the break and ECCS valves is significantly larger than the energy release to the RPV by core heat transfer. The NRELAP5 calculated the heat transfer into the CNV wall results in continuous depressurization of both the RPV and CNV, after the initial pressurization of the CNV.

The NuScale analysis shows that the core remains covered and that MCHFR is not violated; thus, the peak cladding and fuel centerline temperatures are the values at a steady state before the LOCA. Based on its review of the applicant’s representative calculations for liquid space breaks described in this section, the NRC staff concluded that the sample analyses appropriately illustrate that implementation of the methodology as specified in the TR will provide conservative and expected results. Steam Space Break

The NRC staff also reviewed NuScale’s analysis of a steam space break that occurs on the RCS high point vent line. Similar to the RCS injection line break discussed in TR Section 9.1, “Loss-of-Coolant Accident Progression in the NuScale Power Module,” the 100 percent break 66

on the high point vent line causes a reactor trip signal based on the high containment pressure followed by the reactor trip and containment and secondary isolation. NRELAP5 calculates reduced recirculation flow after the reactor trip; and NRELAP5 calculates a brief period of flow reversal in the average core channel with upward flow re-established on both average and hot fuel assemblies during the remainder of the transient.

The NuScale analysis shows that the MCHFR in the transient is at a steady-state and the MCHFR margin increases with time, due to the power and flow mismatch. Based on its review of the applicant’s representative calculations for steam space breaks described in this section, the NRC staff concluded that the sample analyses appropriately illustrate that implementation of the methodology as specified in the TR will provide conservative and expected results.

4.9.2 Break Size

The NRC staff reviewed NuScale’s spectrum of break areas for different break locations given in the LOCA TR. The break areas range from [[

.]]

For all breaks, the break size impacts the timing of events because the break flow rate is proportional to the area for similar upstream conditions. The smaller break sizes produce slower depressurization and lower mass/energy loss rates. For example, events for the 10 percent break size take about 10 times as long as compared to the maximum break size.

The area ratios between break area and maximum break area presented in Table 5-7, “Summary of analyzed break sizes,” of the LOCA TR are used to define “scaled time.” This allows the presentation of computed results for different break areas to be presented on the same plot using “scaled time.” This is important to interpret many of the figures presented in Sections 9.2, “Break Size,” and 9.3. “Decay Heat Removal System Availability,” of the LOCA TR.

NuScale presented the results of a spectrum of breaks in the discharge line, injection line, and high point vent, and pressurizer spray supply. In the calculations audited by the NRC staff, as described in the associated audit report (ML20010D112), NuScale assumes a 10 percent tube plugging as well as a fouling equivalent to 10-4 ft crud thickness on the inner surface, in sensitivity studies.

The NRC staff observes that the limiting break was the 5 percent break with the minimum liquid level above the TAF. This behavior was stated as a “direct” result of the ECCS actuation determined by the IAB release pressures where the ECCS valves do not open until the core collapsed levels are below the final equilibrium value. The NRC staff audited the sensitivity studies applied to this break spectrum, as described in the associated audit report (ML2010D112), including nodalization, time step size, CCFL at the pressurizer baffle plate, ECCS valve parameters (IAB release pressure differential threshold, size/capacity), core power distribution radial peaking assigned to the hot assembly, and CNV initial pool temperatures covering the range 140 ºF (60 ⁰C) down to and including 40 ºF (4 ⁰C). The most significant impact from the sensitivity studies were the DHRS unavailability (not credited in the LOCA analyses), IAB release pressure (low is most limiting for minimum liquid level), and proposed variation in ECCS valve sizes. 67

The NRC staff finds that these analyses and sensitivity studies contain sufficient parameter and break size variations to properly identify the limiting 5 percent injection line break that produces the minimum liquid level above the TAF. The NRC staff finds that the limiting small break injection line LOCA is correctly identified, based on the minimum liquid level above the TAF.

The NRC staff further observes that for the very small breaks where depressurization is very slow, the ECCS valves could remain closed for long periods of time, resulting in large losses of primary liquid. Most important however, is the heat transfer from the RPV to the liquid in the CNV which accumulates the lost primary liquid. It is the conduction of heat from the RPV into the CNV, in addition to the break that depressurizes the RPV to conditions that allow the ECCS valves, via the IAB release set point, to eventually open and thereby assure long term cooling. Since the operation of the DHRS is not credited in the LOCA analyses, the NRC staff considers LOCA events, particularly the smaller breaks in the spectrum, to be conservatively treated since this additional heat removal and depressurization mechanism provided by the DHRS is not credited. As such, the NRC staff notes that the IAB release set point, heat removal from the RPV into the CNV due to the lower temperature liquid in this region, and the DHRS all can effectively contribute to depressurizing the RPV and provide a means of assuring post-LOCA long term cooling for the NuScale NPM.

Based on its review of the applicant’s representative calculations for different break sizes described in this section, the NRC staff concludes that the sample analyses appropriately illustrate that implementation of the methodology as specified in the TR will provide conservative and expected results.

4.9.3 Decay Heat Removal System Availability

The discussion in the previous sections indicates that the applicant took no credit for the DHRS operation. The DHRS adds an additional heat sink capacity during the NPM LOCA that is of benefit to the collapsed level above the TAF, primarily for the smaller break sizes. When the DHRS operation is taken into account, all break sizes behave similarly and minimum CLLs that are maintained very close to the final equilibrium level for most all of the break sizes. The NuScale reported analysis demonstrated that more adverse conditions are not created when crediting the DHRS. The NRC staff finds that not crediting DHRS for LOCA analyses is a conservative approach.

4.9.4 Power Availability

The discussion in the previous sections assumes that both ac and dc power are available during the NPM LOCA. Loss-of-power is defined as either loss of only ac power or loss of both ac and dc power. The loss of all power causes an immediate reactor trip and de-energizes the ECCS valves. However, the ECCS valves are not opened until the differential pressure between the RPV and CNV reach the IAB release set-point.

The NRC staff reviewed NuScale’s analysis results showing that the loss of both ac and dc power has a significant impact on the steam space breaks, down to a two percent break size, but has a minimal impact on the liquid space breaks. The resulting peak containment pressures are not significantly higher than with all power.

Based on its review of the applicant’s representative calculations for different assumptions on power availability described in this section, the NRC staff concludes that the sample analyses 68

appropriately illustrate that implementation of the methodology as specified in the TR will provide conservative and expected results. 4.9.5 Single Failure

The NRC staff reviewed NuScale’s results of analyses assuming the single failures described in Section 9.5, “Single Failure,” of the LOCA TR. As described in Section 4.5.4 of this SER, specific LOCA event limiting single failures are evaluated as part of a design-specific application of this methodology, such as the NuScale DCA. This is reflected in item 6 under Section 6.0 of this SER. Based on its review of the applicant’s representative calculations for different single failure assumptions, the NRC staff concludes that the sample analyses appropriately illustrate that implementation of the methodology as specified in the TR will provide conservative and expected results.

4.9.6 Core Collapsed Liquid Level Calculation

The NRC staff reviewed NuScale’s summary of its calculation of the core CLL, and audited the calculations underlying it as described in the associated audit report (ML20034D464). The NRC staff notes that the minimum core CLL calculation is not the traditional axial formulation but is instead volume based. The axially based approach has some inherent conservativisms since it uses only stacked node height times the computed node liquid fraction, but it is the only method that can capture true “minimum CLL.” The volume-based approach provides a distorted estimate of approximately an additional 2 ft (0.6m) of core level since it credits liquid that is in the riser that is not necessarily available to the core. The volume-based CLL term more appropriately represents “riser volume averaged liquid level” rather than a riser minimum CLL. This method could erode true core CLL margin such that MCHFR may be encountered in the hot channel before the volume-based CLL term reaches the TAF.

Because the LOCA TR acceptance criteria employs the CLL and CHF, both of which must be met, and because the NRC staff finds that the Hench-Levy/Griffith-Zuber correlation employed by the applicant is sufficiently conservative and reasonably bounding in terms of predicting CHF, the NRC staff accepts the applicant’s use of a volume-based CLL. The NRC staff notes that the use of a volume-based CLL in the absence of a CHF criteria would not be acceptable, for the reasons cited above.

4.9.7 Sensitivity Studies

The NRC staff reviewed NuScale’s evaluation of the sensitivity of the LOCA EM results to the changes in modeling parameters summarized in Section 9.6, “Sensitivity Studies,” of the LOCA TR and audited the underlying calculations, as described in the associated audit report (ML20010D112). These parameters included nodalization, time-step size, counter current flow at the pressurizer baffle plate, and ECCS valve parameters (IAB release pressure, differential threshold, valve size/capacity, and valve stroke time). The NRC staff also reviewed NuScale’s evaluation of the sensitivity of the LOCA results to the core power distribution, including axial power shape and hot fuel assembly radial peaking, and initial reactor cooling pool temperature.

The NRC staff also reviewed how NuScale used the sensitivity studies shown in Section 9.6 to support its selection of input values and modeling assumptions for its LOCA EM. These NuScale studies are limited to the range of breaks (2.23 in2 to .05 in2) as defined by the NuScale break spectrum given in Section 5.4, “Loss-of-Coolant Accident Break Spectrum, “of 69

the LOCA TR. Based on its review of the applicant’s representative sensitivity studies for varying the modeling parameters described in this section, the NRC staff concludes that the sample analyses appropriately illustrate that implementation of the methodology as specified in the TR will provide conservative and expected results.

Model Nodalization

The NRC staff reviewed NuScale’s nodalization sensitivity study to determine the impact of nodalization on the key LOCA FOMs including containment pressure and collapsed RPV riser liquid level. To assess the impact of nodalization on the NPM LOCA behavior, NuScale evaluated the three nodalization schemes shown in Table 9-3, “Number of volumes in reactor pressure vessel and containment vessel nodalization,” of the LOCA TR for the full range of break sizes for both the RCS injection line and the high point vent line breaks.

NuScale evaluated both break locations without DHRS operation, loss-of-power, and with no single failure. The NRELAP analyses results for the 100 percent high point vent line break are similar for the different NRELAP5 nodalization schemes. The coarser NRELAP5 nodalization in the containment generates slightly early ECCS actuation signal compared to the coarse and fine nodalization. NRELAP5 calculates similar LOCA response in RPV and CNV pressures and collapsed levels for the high point vent line break scenario.

The NRC staff did not identify any issues with the NuScale selected NRELAP5 nodalization for the calculation of the riser CLL. The NRC staff noted that the accumulation of subcooled water in the CNV resulting from a LOCA blowdown, is important to the computation of the maximum containment pressure. The nodal solution of NRELAP5 must be able to capture the stratification of subcooled water during blowdown for the determination of the maximum CNV pressure.

The NRC staff reviewed information relative to NuScale’s CNV response analysis (ML18298A360 and ML19073A241) that evaluated the effects of using a set of coarser and finer axial nodalizations for the CNV volume, a finer reactor pool nodalization, and a finer CNV heat structure radial nodalization to determine the most limiting nodal representation with respect to CNV peak pressure and temperature. The information that the NRC staff reviewed, showed the expected thermal stratification with minimal temperature differences between the three nodal selections. The NRC staff also reviewed the applicant’s information (ML19151A837) providing similar NRELAP5 containment model nodalization studies for the NIST-1 facility containment for the HP-49 test, which involved the largest RPV liquid discharge into the containment with the lowest elevation. The NRC staff concluded that the NIST-1 nodalization study results and convergence trends are very similar to that of the NPM nodalization studies, and the sensitivity of peak containment pressure to nodalization is not significant. Thus, the NRC staff concluded that the nodal selection has a small impact on the maximum CNV pressure.

Based on the discussion above, the NRC staff found that the NRELAP5 nodalization, used by NuScale, is appropriate to conservatively predict the CLL and CHF margin.

Time-Step Size Selection

NRELAP5 restricts time-step size by the courant time-step size and the accumulation of the mass-error during the time integration. NRELAP5 LOCA simulations set the courant time-step size to evaluate the effect of time-step size selection on the key NPM LOCA FOMs. The NRC

staff reviewed Figures 9-26 through Figures 9-29 in the LOCA TR for the full-size injection line and high point vent line breaks, which illustrate that the maximum time-step size allowed for the NRELAP5 calculations is mainly determined by the mass-error management. These figures show that the containment and RPV pressures, minimum collapsed level above the TAF in the RPV riser, hot channel mass flux, and hot channel MCHFR, are all independent of the time-step sizes selected for the simulation.

Based on the information that the NRC staff reviewed above, and provided that all NRELAP5 calculations continue to show that the collapsed liquid riser level remains above the TAF, as is specified in the LOCA TR, the NRC staff finds the NRELAP5 time step selection process to be acceptable.

Counter Current Flow Limitation Behavior on Pressurizer Baffle Plate

The NRC staff reviewed NuScale’s use of the Bankoff CCFL correlation at the pressurizer baffle plate with a slope of 1.0. A few of the break spectrum cases activated the CCFL flag at the pressurizer baffle plate, which did not allow liquid to readily drain from the pressurizer to the downcomer in the presence of upward steam flow. These break cases were limited to the larger pressurizer spray and vent line breaks. The NRC staff reviewed NuScale’s analysis of liquid and steam breaks to assess the effects of increasing the Bankoff CCFL model slope between 1.0 and 2.0.

The NRC staff noted that the change in the CCFL slope had a significant effect on the immediate NRELAP5 calculated pressurizer level and this change also affected the instantaneous collapsed riser liquid level as shown in Figure 9-30, “Effect of counter current flow limitation line slope on levels for 100 percent high point vent line break,” for the high point vent line break. A higher CCFL slope causes a lower CLL above the TAF because the water is held up in the pressurizer for longer periods of time. However, this change in the Bankoff slope did not impact the riser level for the entire transient because the pressurizer eventually empties and the responses merge before reaching the minimum CLL above the TAF. Hence, the slope input for the CCFL correlation has no impact on the FOM for the minimum collapsed riser level.

Because all NRELAP5 LOCA analyses continue to show that the pressurizer has completely emptied well in advance of the calculated riser CLL reaching the minimum value above the TAF, the NRC staff accepts the Bankoff slope used by NuScale.

Emergency Core Cooling System Valve Parameters

The staff reviewed NuScale’s modeling of the ECCS valve characteristics. The NuScale DCA provides minimum and maximum valve sizes and a range of differential pressures at which the IAB arming valve closes (locks) and opens (releases). The staff reviewed NuScale’s evaluation of liquid and steam breaks, which evaluate separate and combined effects of the range of these valve characteristics on the NRELAP5 calculated LOCA FOMs. The NRELAP5 calculated minimum riser collapsed liquid level shows no dependence on IAB release pressure for the larger breaks. However, for the smaller breaks (less than 35 percent), the staff noted that the minimum collapsed liquid level decreases with decreases in IAB release pressure because the ECCS actuation is determined by the IAB lockout and release pressures where the valves do not open until collapsed levels are well below the stabilized level entering the long-term cooling phase.

The staff also noted that large breaks (greater than 35%) cause relatively rapid RCS depressurization, and CNV pressure is not affected by the IAB release pressure. For the smaller breaks (less than 35 percent area), the peak CNV pressure is higher at a higher IAB release pressure because ECCS activation is determined by the IAB release pressure. However, the peak CNV pressures for these smaller breaks are still lower than the peak CNV pressure for larger breaks.

The NRC staff reviewed NuScale’s sensitivity results, shown in LOCA TR Figures 9-32 and 9- 33, which illustrate the effect of RRV and RVV valve sizes on the peak CNV pressure and the minimum CLL as functions of break size. Overall, the impact of the ECCS valve size on the peak CNV pressure and the collapsed riser liquid level is small. NuScale concludes that a larger RRV size and the lower IAB release pressure set points, generate lower minimum riser collapsed levels above the TAF. The NRC staff agrees with this conclusion as it is expected since larger RRV and delayed IAB release would increase RCS inventory lost before ECCS is actuated.

Additionally, NRC staff performed sensitivity of ECCS results with riser flow holes added and the expanded ECCS actuation signal based on RCS pressure that is interlocked with RCS hot temperature and CNV pressure. This feature results in earlier actuation of ECCS on pressure, particularly for steam space breaks and those with DHRS active. Small liquid space breaks and those without use of DHRS may continue to actuate on high CNV level.

Initial Reactor Pool Temperature

The NRC staff reviewed NuScale’s sensitivity studies covering the range of initial pool temperatures to investigate the impact of the pool temperature on NRELAP5 calculated LOCA EM FOMs. NuScale evaluated the reactor pool temperatures, ranging from 40 ⁰F (4 ⁰C) to 140 ⁰F (60 ⁰C). NuScale evaluated that the RCS injection line breaks down to 5 percent of the full- break size break area, analyzed. The NRC staff noted that the effect of the initial pool temperature on the peak CNV pressure is more pronounced for smaller breaks. For a 100 percent injection line break, the break energy and energy release by the ECCS valve flows are larger than that of the CNV to the reactor pool energy transfer. The energy transfer from the CNV wall to the reactor cooling pool becomes comparable to the break energy for the smaller breaks; this results in significantly lower peak CNV pressures for smaller breaks. However, the maximum peak CNV pressure for all the break sizes considered, occurs at larger breaks. The applicant therefore concluded that the initial pool temperature has negligible impact on the maximum peak CNV pressure as NuScale’s LOCA EM FOM. The slight increase in the peak CNV pressures at larger break sizes is conservatively considered by biasing the initial pool temperature to its maximum value. The applicant also reached a similar conclusion for the minimum CLL above TAF, as the maximum pool temperature produces lower momentary minimum levels at smaller break sizes. For all the initial pool temperatures investigated in the sensitivity calculation, the NRC staff notes that no CHF violation is observed; therefore, the minimum MCHFR is defined by the steady state value. Based on its review of NuScale’s sensitivity studies, the NRC staff agrees that the maximum pool temperature is a conservative assumption for the NuScale maximum CNV pressure analysis.

Core Power Distribution

The NRC staff reviewed NuScale’s sensitivity study of the impact of core power distribution for the full-range of the RCS injection line break sizes including axial power shapes and radial hot 72

fuel assembly power peaking to investigate the effect of core peaking on NRELAP5 calculated LOCA FOMs. NuScale choose axial power shapes to represent [[

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[[ ,]] the NRC staff finds that the core power distribution used in the NuScale LOCA EM, is acceptable. 4.9.8 Loss-of-Coolant Accident Calculation Summary

The NRC staff reviewed NuScale’s summary conclusions based on its break spectrum calculations and sensitivity studies as listed in the LOCA EM TR Section 9.7, “Loss-of-Coolant Accident Calculation Summary.” NuScale LOCA EM TR Section 9.7 presents a range of representative LOCA analyses used by NuScale to demonstrate that NRELAP5 is capable of evaluating NPM LOCAs against the primary FOMs (i.e. maintaining the CLL in the riser above the TAF and maintaining the CHF ratio above the MCHFR).

The NRC staff finds that the accidents evaluated appropriately cover the range of applicability to adequately show that NRELAP5 is capable of performing LOCA analysis to support the NuScale DCA.

In addition, the NRC staff performed extensive confirmatory analyses independently using both the TRACE and NRELAP5 computer codes. The scope of the confirmatory analyses includes the following categories of calculations:

a. TRACE NuScale NPM model development

The NRC staff noted in audit reviews, as documented in the associated audit reports (ML20010D112), that NuScale used an ANSYS solids model to develop all the NPM geometry and mass information based on “cold” RCS dimensions from released drawings. Using the guidelines developed for NIST-1 modeling, NuScale used this NPM design information to develop the base NRELAP5 NPM input model. The NRC staff then used the NPM solids model design information along with the applicant’s NRELAP5 base model to develop the NPM TRACE model, which models the core, the SG, the RPV and the containment with 3-D VESSEL components. TRACE VALVE components, control blocks and signal variables were used to model the ECCS valves and the RCS protection systems.

b. TRACE NPM Best Estimate LOCA and IORV Analysis

Using the developed TRACE NPM model, the NRC staff performed a spectrum of LOCA analyses of different LOCA break locations and sizes. Sensitivity calculations were performed to confirm LOCA progression trends and to evaluate 73

the impact of different single failure assumptions, IAB block and release pressure set points, and investigate margins to the key FOMs.

c. NRELAP NPM LOCA, IORV and Containment Pressure Analysis

The NRC staff audited NuScale NPM computer code and modeling, as described in the associated audit report (ML20010D112), and incorporated the NRELAP5 code and input models into the NRC’s Symbolic Nuclear Analysis Package (SNAP) user interface processor for performing engineering analysis. Using these input models, the NRC staff performed LOCA and containment peak pressure analysis for the limiting break locations and the opening of different ECCS valves to better understand key characteristics of the NPM design.

d. TRACE NIST-1 Benchmark Analysis (HP-02, HP-05, HP-06b, HP-43, and HP-49)

Similar to studies for NPM, the NRC staff used NIST-1 design information and the applicant’s NRELAP5 models to develop the NIST-1 TRACE model which modeled the core, the RPV and the containment with 3-D VESSEL components. Using the developed TRACE models and the audited NIST-1 data, the NRC staff performed independent assessments for the key NIST-1 tests.

e. NRELAP NIST-1 Benchmark Analysis (HP-02, HP-06b, and HP-49)

The NRC staff also performed sensitivity assessments with NRELAP5 and the applicant’s modeling to better understand unique behavior and characteristics and results of the NIST-1 modeling and test benchmark results.

All these benchmarks and calculations confirmed to the NRC staff that the LOCA EM has adequate basis and validation to sufficiently predict key FOMs for the NPM design (i.e., the collapsed water level remains above the TAF, the MCHFR is not violated, and the peak containment pressure is much lower than the design limit pressure).

5.0 EVALUATION MODEL FOR INADVERTANT OPENING OF RPV VALVES

5.1 Event Description and Classification

An accidental IORV (i.e., RSV, RVV, or RRV) results in reactor vessel depressurization and a decrease of reactor vessel coolant inventory that could be caused by a spurious electrical signal, hardware malfunction, or operator error. The EM and methodology applied to analyze SRP Section 15.6.1, “Inadvertent Opening of a PWR Pressurizer Pressure Relief Valve,” and SRP Section 15.6.6, “Inadvertent Operation of the Emergency Core Cooling System (ECCS),” events are developed by extending the LOCA Methodology. These inadvertent RPV valve events are classified by NuScale as AOOs.

The applicant’s description confirms that each ECCS valve system includes an IAB valve to block the main ECCS valve from opening based on a set release pressure difference threshold between the CNV and RPV. The IAB effectively reduces the frequency of inadvertent openings

of the valves during power operation. The limiting event analyzed is the mechanical failure of an ECCS valve that depressurizes the control chamber at operating pressure.

5.2 Evaluation Model

The NRELAP5 model utilized for IORV analysis is developed from a base plant model that is modified for the important aspects for the IORV, with similar modifications that are applied to the LOCA EM as described in Section 5.1, “NRELAP5 Loss-of-Coolant Accident Model for the NuScale Power Module,” of the LOCA TR. The NRC staff reviewed the primary differences in the modeling, which are related to those necessary to better align the analysis with AOO acceptance criteria instead of postulated accident criteria.

The NRC staff reviewed the overall EM objectives for mitigation, which are the same as for LOCAs in that: (1) the CNV must contain the loss of inventory from the RCS, (2) the remaining ECCS valves must actuate to depressurize the RPV into the CNV until pressure equalization, which allows the return of discharged fluid back into the RPV to cool the core, and (3) stable natural circulation flow must be maintained via ECCS steam condensation cooling to the reactor pool. Initial plant conditions are conservatively biased similar to the initial conditions used for LOCA analyses.

The applicant used NRELAP5 LOCA modeling methods for its analysis of IORV events because the transient progression of the event and the PIRT phenomena are similar to that of LOCA pipe breaks. Therefore, the NRC staff’s review focused in the areas of the differences from the LOCA modeling. The method specifies selection of input parameters and initial conditions to provide a conservative calculation relative to MCHFR since IORV is the limiting event for CHFR. The NRC staff reviewed and audited the core modeling, as described in the associated audit report (ML20034D464), that is based on heat transfer options that specify the extended Hench- Levy CHF correlation for high flow conditions, and the Griffith-Zuber CHF correlation for low flow conditions. The initial conditions and biasing for the steady state portion of the transient are the same as for a LOCA.

The NRC staff reviewed the information on those differences (ML18264A338), which are:

1. The ECCS valve opening stroke is reduced to 0.1 seconds for faster opening, which produces higher flow rates and faster depressurization.

2. The ECCS valve inadvertent opening choked flow is modeled as a break with Moody/Henry-Fauske (c=3) to maximize two-phase flow which is consistent with LOCA break methodology.

3. An Additional 2 second scram delay conservatism used in a LOCA is removed (not used for AOO calculations).

4. The 15 percent bias on fuel thermal conductivity and heat capacity to increase stored thermal energy for a LOCA is removed (not used for AOO calculations).

5. A fuel gap conductance is varied via sensitivity analyses instead of the bounding value used in a LOCA.

6. Although the same CHF correlation methods are used, less limiting 95/95 tolerance limit of 1.13 is applied in AOO calculations. The NRC staff further reviewed the following modifications and audited the associated underlying calculations and sensitivity analyses, as described in the associated audit reports (ML20034D464). The initial conditions and biasing included sensitivity cases for: (1) all electric power available, (2) loss of normal ac power, and (3) loss of normal ac and dc power. The loss of dc power will impact the LOCA progression by immediately triggering the ECCS valves to go to their fail-safe position, where each valve is held closed by its IAB.

The single failure assumptions included: (1) no single failure, (2) failure of a single RVV to open, (3) failure of a single RRV to open, and (4) failure of one ECCS division (i.e., one RVV and one RRV). The NRC staff noted that the failure of an IAB to block (which would cause a second ECCS valve to prematurely open above the IAB threshold release pressure) was not considered. The treatment of the IAB valve’s function to close relative to single failure is discussed further in Section 4.5.4 of this SER. Therefore, for analysis of these events, based on the applicant’s submittal and the discussion in Section 4.5.4 of this SER, the inadvertent operation of ECCS consists of the opening of one RVV or one RRV.

The NRC staff also reviewed and audited sensitivity cases on model parameters for fuel rod gap conductance, axial power shape, ECCS valve sizing, ECCS valve opening rate, and DHRS availability. The IORV event is modeled as a mechanical failure resulting in opening of a single ECCS valve. The valve opening is the initiating event and partial valve opening is not considered as a single failure.

Specific limiting single failures, electrical power assumptions, and whether any operator actions are needed for this event are evaluated as part of a design-specific application of this methodology, such as the NuScale DCA. This is reflected in item 6 under Section 6.0 of this SER.

For the typical inadvertent RVV or RRV opening with or without the loss of ac power, the MCHFR occurs very early in the transient before the rods are fully inserted from the reactor trip on high containment pressure. The remaining ECCS valves open much later, when the ECCS system actuates on the high containment level. Minimum water level above the core occurs as the RPV and containment water levels equalize. The overall RRV transient is like the RVV. However, the liquid-space discharge results in a slower depressurization, accompanied by a greater decrease in core inlet flow as coolant discharges from the downcomer region into the containment. The liquid-space discharge generates an ECCS actuation signal on the high CNV level that occurs earlier than for the RVV transient. After the remaining ECCS valves open, the RRV scenario and the RVV scenario follow similar trends for fluid conditions and heat transfer for long term cooling.

The NRC staff reviewed NuScale’s sensitivity analyses for key model parameters including: (1) Fuel Rod Gap Conductance, (2) Axial Power Shape, (3) ECCS Valve Sizing, (4) ECCS Valve Opening Stroke Time, (5) DHRS Operation, (6) Single Active Failures, and (7) Electric Power Availability. The NRC staff also reviewed the limiting MCHFR case for RVV inadvertent opening with power available, high RCS average temperature, low RCS flow, high RCS pressure, maximum gap conductance, and high PZR level, however the NRC staff’s review of the final limiting case will be documented in the NRC staff’s evaluation of a NPM application (e.g. the staff’s SER for the NuScale DCA). The NRC staff’s review was intended to confirm that the

LOCA/IORV methodology is capable of predicting the limiting cases, and the expected key event sequences and parameter trends.

The applicant’s assessment methodology for the IORV events is based on LOCA experimental test benchmarks, inclusive of SETs for CHF and IETs for phenomenological system responses (per Table 7-1, “NRELAP5 loss-of-coolant accident assessment matrix,” of the LOCA TR). The high-ranked phenomena derived in the NuScale LOCA PIRT are indicated by the applicant as identical to the IORV event, since the progression of the transient is very similar. The NRC staff noted that although the progression is similar, the effective break flow areas of IORV events are much larger. The LOCA break sizes were limited to 2.0-inch pipes in CVCS or PZR spray, whereas limiting flow areas for ECCS valves are much larger, 2.5-inch for the RRVs and 4.75- inch for the RVVs. Additionally, the NRC staff observed that the RRVs are located at a lower elevation relative to the TAF in comparison to the CVCS charging and discharge lines. The LOCA PIRT high-ranked phenomena (per Table 4-4, “High-ranked phenomena,” of the LOCA TR) included and evaluated the RRVs and RVVs but only as related to ECCS actuation, i.e., phase 1b, and not as an initiating event.

The NRC staff concludes that since the RRV flow area and location within the RPV are relatively similar, two-phase critical flow phenomena (knowledge level 2) can be extended from the CVCS break assessment. For the RVVs, the flow area is significantly larger but the critical break flow phenomena is primarily single phase (knowledge level 4).

The applicant additionally uses the NIST-1 Integral Effects Tests, HP-06b and HP-43, to provide assessments for NRELAP5 predictions of key critical flow phenomena. The HP-43 test is an updated version of the HP-09 test where the depressurization of the RPV is initiated by a spuriously opened RVV sized to represent one of three RVVs, and the remaining RVV, which opens on ECCS actuation, is sized to represent flow from two RVVs. The NRC staff audited this test and the test assessments, as described in the associated audit report (ML20034D464), and concludes that the code predictions adequately matched the data trends and key FOMs. The NRC staff also reviewed supplementary information (ML18268A365) and audited the applicant’s test results and assessment reports for NIST-1 HP-49, as described in the associated audit report (ML20034D464), which provide data for a spurious RRV opening event. The NRC staff similarly concluded that these code predictions also adequately matched the FOMs.

Based on its review of the applicant’s representative calculations for the different IORV events with single failure assumptions described in this section, the NRC staff concludes that the sample analyses appropriately illustrate that implementation of the IORV methodology as specified in Appendix B of the TR (1) is consistent with LOCA EM methodology such that the validations for LOCA are applicable to the IORV event and (2) will provide conservative and expected results.

The NRC staff also performed sensitivity calculations of the IORV results with riser flow holes added and the expanded ECCS actuation signal based on RCS pressure. The results indicated that the limiting CHF are not affected by these design changes.

5.3 Accident Scenario Identification Process

The EM review criterion per RG 1.203, recommends that applicants follow a structured process for the identification and ranking of physical phenomena relevant to the accident scenarios to 77

which the EM will be applied. For the NPM, the applicant has indicated there are no significant differences in the physics of the fundamental phenomena between the LOCA and IORV events, i.e., the initiating locations and effective break flow areas are different but the governing thermal hydraulic code processes would be very similar. NuScale indicated that the high-ranked phenomena from the LOCA PIRT shown in LOCA TR Table 4-4 also apply to the IORV event scenarios, and that the event phases of initial blowdown (1a) and ECCS actuation (1b) are also identical for LOCA and IORV.

5.4 CHF Evaluation

The NRC staff’s evaluation of the high-flow and low-flow CHF models uses the critical boiling transition model assessment framework developed in Appendix A of the safety evaluation (SE) for TR-0116-21012-A, “NuScale Power Critical Heat Flux Correlations,” (ADAMS Accession No. ML18360A632). This framework assesses the CHF correlation through a top-down approach whereby the high level finding that, “the CHF correlation is acceptable,” is broken down into lower level goals (G). The lowest level goals are directly supported by evidence.

In addition to the material submitted by NuScale, the NRC staff considered the historical and publicly available data from NUREG/CR-1559, “Transient Critical Heat Flux and Blowdown Heat-Transfer Studies,” (Office of Science and Technical Information (OSTI) Identifier 5824873) during their review.

The NRC staff notes that information pertaining to the high-flow and low-flow CHF models is also presented in Section 5.1.8.3, Section 6.11, and Section 7.3 of the TR. The NRC staff’s review of the CHF information in TR-0516-49422 is provided in Section 5.5 below.

5.5 Experiential Data

Section B.5.1 of the TR describes the testing used to validate the high-flow CHF correlation, which was obtained at the KArlstein Thermal HYdraulic test loop (KATHY) in Karlstein, Germany. Additionally, Section 6.11, Section 7.3, and Section 5.3 of the TR clarify that Stern CHF data is also used in the development of the high-flow and low-flow CHF correlations as implemented into NRELAP5. The NRC staff previously reviewed the experimental data supporting the CHF model development as part of the review of TR-0116-21012-A, “NuScale Power Critical Heat Flux Correlations,” (ADAMS Accession No. ML18360A632). Accordingly, the NRC staff’s findings for all of the items under G1 of the assessment framework, documented in the SE for TR-0116-21012-A, are applicable to the evaluation of the high-flow and low-flow CHF correlations with the exception of G1.3.1 and G1.3.2. G1.3.1 and G1.3.2, which are evaluated for the high-flow and low-flow CHF correlations below.

Equivalent Geometries

The test bundle used in the experiment should have equivalent geometric dimensions to that of the fuel bundle used in the reactor for all major components.

G1.3.1, Review Framework for Critical Boiling Transition Models

The NRC staff’s SE for TR-0116-21012-A, “NuScale Power Critical Heat Flux Correlations,” (ADAMS Accession No. ML18360A632) established the finding that the test bundles used to obtain validation data for the NSP2 and NSP4 CHF correlations have equivalent dimensions to that of the fuel bundle in the reactor for all major components because: (1) the data collected from the KATHY test loop used test bundles that are representative of prototypical NuFuel- HTP2TM fuel, and (2) the KATHY test loop data was used for the validation of the NSP2 and NSP4 CHF correlations. This finding is applicable to the high-flow CHF correlation because the KATHY data is used to validate this correlation. However, the Stern data is used to validate the low-flow CHF correlation, and the test bundle used at Stern was a preliminary prototype fuel that is not prototypical of NuFuel-HTP2TM fuel.

Section B.5.3 of the TR describes that no KATHY test data exists in the range where the low- flow CHF correlation is used. Therefore, the applicant proposed to use the subset of Stern data that was collected for lower flow rates. In Section 6.1, “Comparison of Stern Preliminary Prototypic to KATHY NuFuel-HTP2™ Test Data,” of TR-0116-21012-A, “NuScale Power Critical Heat Flux Correlations,” (ADAMS Accession No. ML18360A632), the applicant compared the results of CHF testing from Stern to the results from the KATHY test loop. This comparison demonstrated that [[

]] Based on the information provided in Section 6.1 of TR-0115- 21012-A, the NRC staff finds the use of Stern data, to validate the low-flow CHF correlation, to be acceptable because the applicant demonstrated that the grid spacers used in the NuFuel-HTP2TM fuel design do not adversely impact CHF performance.

Equivalent Grid Spacers

The grid spacers used in the test bundle should be prototypical of the grid spacers used in the reactor assembly.

G1.3.2, Review Framework for Critical Boiling Transition Models

The same argument that was applied to G1.3.1 also applies to G1.3.2. Specifically: (1) the KATHY test loop data that is used to validate the high-flow CHF correlation used grid spacers that are prototypical of the grid spacers used in the reactor assembly, and (2) the Stern test data that is used to validate the low-flow CHF correlation does not use a prototypical grid spacer but is acceptable because the applicant demonstrated that the grid spacers used in the NuFuel- HTP2TM fuel design do not adversely impact CHF performance.

5.5.1 Model Generation

5.5.1.1 Mathematical Form

Sections 6.11.3, [[ ]] and 6.11.4, [[ ]] of the TR describe the high-flow and low-flow CHF correlations, respectively. NuScale selected historical CHF models as the basis for CHF modeling in the TR.

The NRC staff compared the high-flow CHF correlation with the [[ ]] presented in Todreas and Kazimi5 and finds them to be consistent. However, NuScale [[ ]] (discussed further in Section 5.5.1.2 of this SER). The NRC staff compared the low-flow CHF correlation with the [[ ]] presented in Appendix G of NUREG/CR-1559 and finds that the parameters and forms are consistent. The NRC staff observed that the coefficient for the low-flow CHF model, given by Equation 6-105 of the TR, differs from the value provided in NUREG/CR-1559 but it is acceptable because the coefficient used in the TR is smaller, resulting in a conservative calculated value for CHF. Additionally, the low-flow CHF correlation uses a CHF multiplier which, as discussed further in Section 5.5.2 and 5.5.3.2 of this SER, resulted in the NRC staff creating Limitation 7 on the application of the low-flow CHF correlation. Based on the information provided in Sections 6.11.3 and 6.11.4 of the TR, and subject to Limitations 7 and 8 for correlation applicability ranges, the NRC staff finds that the high-flow and low-flow CHF correlations contain the appropriate parameters (G2.1.1) and have an acceptable model form (G2.1.2) because these models are consistent with historical models that have been tested and validated.

5.5.1.2 Model Coefficient Generation

5.5.1.2.1 Training Data

Training Data

The training data (i.e., the data used to generate the coefficients of the model) should be identified.

G2.2.1, Review Framework for Critical Boiling Transition Models

Sections 6.11.3 and 6.11.4 of the TR describe the high-flow and low-flow CHF correlations as historical correlations with a modification to the high-flow CHF correlation. Section 6.11.3 of the TR describes the modification to the high-flow CHF correlation as a [[ ]]. Specifically, the discussion clarifies that [[ ]] using CHF data collected at Stern Laboratories because the data collected at Stern Laboratories [[ ]]. Section B.5.3 of the TR discusses the development of the CHFR limit (i.e., validation) for the high-flow CHF correlation as using data collected at the KATHY test loop for prototypical Nufuel-HTP2TM fuel, which is independent of the data collected at Stern Laboratories. Section 6.11.4 of the TR clarifies that no data was collected to train the low-flow CHF correlation (i.e., the collected data was used only for validation purposes). Based on the information provided in Section 6.11.3, 6.11.4, and B.5.3 of the TR, the NRC staff finds the selected training data to be acceptable because it is independent of the validation data.

5 N. E. Todreas and M.S. Kazimi (1990), Nuclear Systems I – Thermal Hydraulic Fundamentals, Taylor & Francis 80

5.5.1.2.2 Coefficient Generation

Coefficient Generation

The method for calculating the model’s coefficients should be described.

G2.2.2, Review Framework for Critical Boiling Transition Models

Section 6.11.3 of the TR describes the curve fitting for the [[ ]] of the high-flow CHF correlation. The resulting coefficients and [[ ]] are provided in Table 6-5 and Figure 6-6 of the TR, respectively. Based on the description provided in Section 6.11.3 of the TR, and the results shown in Figure 6-6 of the TR, the NRC staff finds the coefficient generation to be acceptable because the resulting [[ ]] trends are consistent with the data from Stern Laboratories.

5.5.2 Model Validation

5.5.2.1 Validation Error

Validation Error

The validation error has been correctly calculated.

G3.1, Review Framework for Critical Boiling Transition Models

NuScale uses a database to assess and validate the high-flow and low-flow CHF correlations, which are based on historical CHF correlations. The validation data for the high-flow and low- flow CHF correlations are obtained from testing the NuFuel-HTP2TM simulated fuel. The high- flow correlation applies the [[

]]

Section B.5.3 of the TR explains that a comparison between the NRELAP5 predicted CHF values and measured KATHY test values [[

.]] The NRC staff finds this approach to be acceptable because it results in a conservative CHFR limit.

Section B.5.3 of the TR explains that there is no KATHY CHF test data for the low-flow CHF correlation conditions. Therefore, the CHF limit derived from the Stern CHF data described in TR Section 7.3.6, “Assessment Results,” is used. [[

.]] Therefore, the NRC staff finds this approach to be acceptable because it results in a reasonable CHFR limit and the CHFR limit is typically challenged within the range of the high-flow correlation.

Section 7.3.5 and B.5.2 of the TR describes the NRELAP5 model for the CHF tests. The TR states that the [[

]]. The NRC staff reviewed the NRELAP5 test models and comparisons and confirmed that an acceptable measured value is used for the comparison.

Based on the acceptable comparisons of measured (test data) vs predicted (calculated) power, the NRC staff finds that the calculated validation error (CHFR Limits) is acceptable and is calculated using an acceptable measured value.

5.5.2.2 Data Distribution

5.5.2.2.1 Validation Data

Validation Data

The validation data (i.e., the data used to quantify the model’s error) should be identified.

G3.2.1, Review Framework for Critical Boiling Transition Models

TR Sections 7.3.6 and B.5 describes the validation process used to establish the CHF correlation limits for the high-flow and low-flow CHF correlations. The methodology for determining the correlation limit is identical to that used in the development of the NSP2 and NSP4 CHF correlations described in TR-0116-21012, “NuScale Power Critical Heat Flux Correlations,” Revision 1, issued November 2017 (ADAMS Accession No. ML17335A089), which the NRC staff has reviewed and approved (ADAMS Accession No. ML18214A480). The referenced TR explains that [[

]]. The limits for these correlations were established by comparison to test data which is completely independent of the historical correlation development. Based on the use of historical correlations for comparison to test data and the [[ ]], the NRC staff finds the use of the data to perform validation and to determine the CHFR limit to be acceptable. 82

5.5.2.2.2 Application Domain

Application Domain

The application domain of the model should be mathematically defined.

G3.2.2, Review Framework for Critical Boiling Transition Models

Table 6-6 of the TR provides the application domain for the high-flow CHF correlation. The low- flow correlation application domain is characterized by [[ ]] in TR Section 6.11.4. [[

]] Therefore, the NRC staff finds the [[ ]] to be acceptable.

The NRC staff confirmed that the test data used to validate the high-flow correlation covers the high-flow correlation application range. The NRC staff observed that the test data used to validate the low-flow correlation [[

]].

The NRC staff finds the definition of the application domain to be acceptable because: (1) the high-flow correlation application domain is supported by test data (2) the low-flow correlation application domain is [[

]].

5.5.2.2.3 Expected Domain

Expected Domain

The expected domain of the model should be understood.

G3.2.3, Review Framework for Critical Boiling Transition Models

In TR-1113-5374-NP, “Critical Heat Flux Test Program Technical Report,” January 2014 (ADAMS Accession No. ML14024A455), Section 2.5, the applicant performed preliminary analyses of the NPM to develop conditions covering normal operation, AOOs and postulated accidents. The NRC staff compared the range of test conditions, identified in Table 2-5, “Range of test conditions,” of TR-1113-5374-NP, with the range of applicability of the high-flow and low- flow CHF correlations, provided in Table 6-6 and Section 6.11.4 of the TR, and determined that the range of test conditions encompasses the range of applicability for the high-flow correlation and is reasonable for the low-flow correlation. Based on the performance studies of the NPM presented in TR-1113-5374-NP, the NRC staff finds the expected range to be acceptable.

5.5.2.3 Data Density

Data Density

There should be an appropriate data density throughout the expected domain.

G3.2.4, Review Framework for Critical Boiling Transition Models

Table 6-6 and Section 6.11.4 of the TR mathematically define the application domains for the high-flow and low-flow CHF correlations, respectively. Based on prior experience with CHF correlation reviews, the NRC staff recognizes that the defined application domain of a CHF correlation contains regions where there are no or sparse underlying experimental data and where the correlation may not be used. The NRC staff conducted an audit as part of the review, which included the data density throughout the expected domain, as described in the associated audit report (ADAMS Accession No. ML20034D464). During the audit, the NRC staff observed several plots that show the data collection within the expected domain. These plots show that the [[

]]. Based on the information observed during the audit, the NRC finds that there is appropriate data density throughout the expected domain.

5.5.2.4 Sparse Regions

Sparse Regions

Sparse regions (i.e., regions of low data density) in the expected domain should be identified and justified to be appropriate.

G3.2.5, Review Framework for Critical Boiling Transition Models

As described in Section 5.5.2.3 of this SER, the expected mass flux and pressure regions have appropriate data density throughout the expected domain. Additionally, Section 5.5.2.3 states that the plot comparisons to data show [[ ]]. Accordingly, the NRC staff finds that the sparse regions in the expected domain are identified and appropriately justified.

5.5.2.5 Restricted Domain

Restricted Domain

The model should be restricted to its application domain.

G3.2.6, Review Framework for Critical Boiling Transition Models

Table 6-6 and Section 6.11.4 of the TR mathematically define the application domains for the high-flow and low-flow CHF correlations, respectively. Additionally, Section 1.1 of the TR states that approval of the EM, which includes the CHF correlations range of applicability, for design basis LOCA events is requested. Section 1.1 of the TR extends the purpose of the EM to include IORV. The NRC staff finds that this is consistent with the established precedent and is sufficient for restricting the domain of applicability in conjunction with the limitations and conditions described in this SER.

5.5.3 Consistent Model Error

5.5.3.1 Poolability

Poolability

The validation error should be investigated to ensure that it does not contain any subgroups that are obviously not from the same population (i.e., non-poolable).

G3.3.1, Review Framework for Critical Boiling Transition Models

Section B.5.3 of the TR provides the process for determining the CHF correlation limits. The NRC staff observed that the NuScale process for developing the CHFR limit involved several statistical tests to determine: (1) whether subregions can be combined (i.e., pooled), and (2) whether the data can be treated as normally distributed. Figure B-4, “CHF Statistical Methods Flow Chart,” of the TR provides the NuScale process for developing the CHFR limit. The statistical tests used by NuScale, and as shown in Figure B-4, are consistent with those described in NUREG-1475, Revision 1, “Applying Statistics,” March 2011 (ML11102A076). Based on the description of these methods in NUREG-1475, the NRC staff finds that these statistical tests are consistent with established guidance and are therefore, acceptable. Additionally, based on NuScale selecting a bounding CHFR limit the NRC staff finds the process used to select the CHFR limit acceptable.

5.5.3.2 Nonconservative Subregions

Nonconservative Subregions

The expected domain should be investigated to determine if it contains any non-conservative subregions which would impact the predictive capability of the model.

G3.3.2, Review Framework for Critical Boiling Transition Models

The NRC staff analyzed the measured-to-predicted performance of the high-flow CHF correlation. The NRC staff audited the subregions used to determine the correlation CHF limit, as described in the associated audit report (ADAMS Accession No. ML20034D464). During the audit, the NRC staff confirmed that the [[ ]]. The NRC staff finds that the high-flow CHFR limit is suitably conservative over the application domain.

The NRC staff analyzed the measured-to-predicted performance of the low-flow CHF correlation. The NRC audited the subregions used to determine the correlation CHF limit, as described in the associated audit report (ADAMS Accession No. ML20034D464). The NRC staff observed that there is a [[

]] discussed in Section 6.11.4 of the TR and the information reviewed during the audit, the NRC staff finds that the low-flow CHFR limit is suitably conservative over the application domain established in the TR.

5.5.3.3 Model Trends

Model Trends

The model is trending as expected in each of the various model parameters.

G3.3.3, Review Framework for Critical Boiling Transition Models

The NRC staff analyzed the measured-to-predicted performance of the high-flow CHF correlation. The NRC staff audited the model trends for measured-to-predicted performance with respect to mass flux, pressure, subcooling and critical quality, as described in the associated audit report (ADAMS Accession No. ML20034D464). During the audit, the NRC staff confirmed that [[

]], the NRC staff finds that the high-flow correlation trends are as expected in each of the various model parameters. 86

The NRC staff analyzed the measured-to-predicted performance of the low-flow CHF correlation. The NRC audited the model trends for measured-to-predicted performance with respect to mass flux, pressure, subcooling and critical quality, as discussed in the associated audit report (ADAMS Accession No. ML20034D464). During the audit, the NRC staff observed that [[

]], the NRC staff finds that the low-flow correlation trends are as expected in each of the various model parameters.

5.5.4 Quantified Model Error

5.5.4.1 Error Data Base

Error Data Base

The model’s error should be calculated from an appropriate data base.

G3.4.1, Review Framework for Critical Boiling Transition Models

As described in Section 5.5.2.2.1 of this SER, NuScale [[ ]] to establish the CHFR limit for the high-flow and low-flow correlations. Based on the use of [[ ]] to set the CHFR limits, the NRC staff finds that the error database used to determine the CHFR limits for the high-flow and low- flow CHF correlations are acceptable.

5.5.4.2 Statistical Method

Statistical Method

The model’s error should be calculated using an appropriate statistical method.

G3.4.2, Review Framework for Critical Boiling Transition Models

Section B.5.3 of the TR describes the statistics used to develop the CHF correlation limits. As evaluated in Section 5.5.3.1 of this SER, the statistical tests for poolability and normality are consistent with established guidance and are therefore acceptable. If normality testing determines that the data are normally distributed, NuScale determines the CHFR limit by adding the standard deviation times an appropriate tolerance factor (i.e., sufficient to establish a 95/95 CHFR limit) to the predicted-to-mean distribution average. If normality testing determines that the data are not normally distributed, then NuScale uses nonparametric statistics to establish a

95/95 CHFR limit. The NRC staff finds NuScale’s statistical methodology acceptable because it is consistent with established guidance.

5.5.4.3 Appropriate Bias for Model Uncertainty

Appropriate Bias

The model’s error should be appropriately biased in generating the model uncertainty.

G3.4.3, Review Framework for Critical Boiling Transition Models

Section 7.3.6 and B.5.3 of the TR develop the CHFR limits for the high-flow and low-flow CHF correlations. The CHFR limit of 1.05 for the high-flow correlation for IORV events, presented in Section B.5.3 of the TR, [[ ]]. NuScale selected a CHFR limit of 1.29 for the low-flow correlation for IORV events, [[ ]]. Additionally, a 3 percent engineering uncertainty factor and a 3 percent fuel rod bowing factor is applied to the CHFR limits for IORV events resulting in final CHFR limits that are rounded up to 1.13 and 1.37 respectively. NuScale selected a CHFR limit of 1.29 for the high-flow and low-flow correlations for LOCA events, [[ ]]. Based on the use of bounding values for the CHFR limit, the NRC staff finds that the CHFR limits for the high-flow and low-flow correlations are appropriately biased. To ensure that the high-flow and low-flow CHFR limits are used in a manner consistent with their approved biases, the NRC staff established Condition 9 on the application of the CHFR limits for IORV and LOCA events.

5.5.5 Model Implementation

5.5.5.1 Same Computer Code

Same Computer Code

The model has been implemented in the same computer code that was used to generate the validation data.

G3.5.1, Review Framework for Critical Boiling Transition Models

Section 5.0 and B.6 of the TR provide the description of the EM which includes the use of the NRELAP5 computer code. Section 7.3 and B.5.3 describes the validation of the computer code to the test data. To ensure that the high-flow and low-flow CHF correlations are used in a manner consistent with their validation, the NRC staff established Limitations 7 and 8 on the use of NRELAP5 calculations using the high-flow and low-flow CHF correlations. Based on the description in Sections 5.0, B.6, 7.3 and B.5.3 of the TR, and pursuant to Limitations 7 and 8, the NRC staff finds that the high-flow and low-flow CHF correlations are implemented using the same computer code used to generate validation data, and are therefore acceptable.

88 5.5.5.2 Same Methodology

Same Methodology

The model’s prediction of critical boiling transition is being applied in the same manner as it was when predicting the validation data set.

G3.5.2, Review Framework for Critical Boiling Transition Models

As described in Section 5.5.5.1, “Same Computer Code,” of this SER, the NRC staff established Limitations 7 and 8 to ensure that the high-flow and low-flow CHF correlations are used in a manner that is consistent with their validation. Based on the description in Sections 5.0, B.6, 7.3 and B.5.3 of the TR, and pursuant to Limitations 7 and 8, the NRC staff finds that the high- flow and low-flow CHF correlations are being applied in the same manner as when predicting the validation data set, and are therefore acceptable.

5.5.5.3 Transient Behavior

Prediction of Transient Behavior

The model results in an accurate or conservative prediction when it is used to predict transient behavior.

G3.5.3, Review Framework for Critical Boiling Transition Models

During an audit with NuScale during June 13 – 15, 2017, as described in the associated audit report (ML17278A168), the NRC staff observed that NuScale performed [[

]] Based on the results of these tests, the NRC staff finds that the high- flow and low-flow CHF correlations provide suitably conservative predictions for CHF when used to predict transient behavior.

5.6 Summary

The NRC staff approves the CHF modeling described in TR-0516-49422, Revision 2, subject to the limitations and conditions identified in Section 6.0 of this SER. In particular, the NRC staff finds that: (1) the high-flow CHF correlation is acceptable for use in performing safety analyses of the NPM with NuFuelHTP2TM fuel, with a CHFR limit of 1.13 for IORV events and 1.29 for LOCA events, over the range of applicability provided in Table 5.2-1, “High-Flow CHF Correlation Range of Applicability,” of this SER, and (2) the low-flow CHF correlation is acceptable for use in performing safety analyses of the NPM with NuFuelHTP2TM fuel, with a CHFR limit of 1.29 for LOCA events and 1.37 for IORV events. These conclusions are based on the following three findings:

89 1. The experimental data supporting the high-flow and low-flow CHF correlations are appropriate as evidenced by meeting all the supporting goals discussed in Section 5.5 of this SER.

2. The high-flow and low-flow CHF correlations were generated in a logical fashion as evidenced by meeting all the supporting goals discussed in Section 5.5.1 of this SER.

3. The high-flow and low-flow CHF correlations have sufficient validation, demonstrated through appropriate quantification of their error, as evidenced by meeting all the supporting goals discussed in Sections 5.5.2 through 5.5.5 of this SER.

6.0 LIMITATIONS AND CONDITIONS

This section provides a summary of the limitations and conditions based on the technical evaluation of the NuScale TR 0516-49422, “Loss-of-Coolant Accident Evaluation Methodology”, Revision 2. As a result of its in-depth technical evaluation, the NRC staff determined that the NuScale LOCA EM, including the methodology to analyze events initiating from the IORV specified in Appendix B, can be used for the NuScale NPM design, subject to the limitations and specific restrictions on the use of this model as listed below.

1. Regulatory compliance with 10 CFR Part 50, Appendix K for application of the LOCA EM for features not evaluated in the TR. An applicant or licensee referencing this report will be required to address regulatory compliance with 10 CFR 50.46 and 10 CFR Part 50, Appendix K, which could include seeking an exemption from the required features not addressed by this EM as described in Table 2-2 of this TR, including: those related to post-CHF heat transfer models; fuel pin models that incorporate clad swelling, rupture and, oxidation; calculation of the metal-water reaction rate using the Baker-Just Correlation and radiation heat transfer.

2. CLL, CHF and Peak Containment Pressure and Temperature Requirements. The NuScale LOCA EM is limited to the evaluation of LOCAs where: (1) the CHF is not exceeded; (2) the CLL remains above the top elevation of the core active fuel region for the full spectrum of break sizes and locations and (3) the containment peak temperature and pressure remain below the design limits.

NRELAP5 does not apply to LOCA conditions where CHF is achieved and core uncovery is predicted to occur since the NuScale LOCA EM has not been demonstrated as adequate to evaluate peak cladding temperature, core wide oxidation, rod swelling and rupture behavior that could occur if the CLL drops below the top of the reactor core sufficiently to cause the active fuel to be exposed to steam cooling. Further, the NuScale design is subject to a potential return to power subsequent to a LOCA with one rod stuck full out of the core. The LOCA EM does not include a method for calculating the number of failed pins from either an initial LOCA or a subsequent return to power with pre-existing fuel pin failures from the LOCA event.

3. No Credit for DHRS Heat Removal in LOCA EM without Further NRC Review and Approval. The NRC staff did not review the DHRS heat removal modeling as part of this overall methodology. Any future credit for DHRS requires review and approval by the NRC.

4. Types of Analyses Approved for LOCA EM.

Use of the LOCA EM is limited to evaluations of the analyses for the FOMs described in the TR: the short term LOCA or an IORV event. The LOCA EM is not approved for use in evaluations for thermal hydraulic analyses not described in the methodology presented in the TR. Use of the LOCA EM is not approved for use in analysis of thermal hydraulic instabilities in the secondary or primary system, peak containment pressure (such as following an IORV), control rod ejection accidents, radiological consequences, non-LOCA events (other than an IORV), return to power analysis assuming the worst- case stuck control rod, and evaluation of the long-term cooling phase.

5. Limitations on NRELAP5 and NPM Model Approval. Use of NRELAP5 is limited to v1.4, in conjunction with NPM model Revision 3, unless changes are made pursuant to a change process specifically approved by the NRC staff for changes to NRELAP5 and the NPM model. NRELAP5 v1.4 and NPM model Revision 3 are approved for use in this TR as part of the LOCA EM. NRELAP5 is not approved for analysis of thermal hydraulic instabilities in the secondary or primary system. When NRELAP5 v1.4 and NPM model Revision 3, as described in this TR, are referenced in other EMs, those applications for use of NRELAP5 v1.4 and NPM model Revision 3 within another EM; require separate approvals to ensure the models and assumptions are defined appropriately for the analyzed FOMs. Use of the NRELAP5 v1.4 and NPM model Revision 3 are therefore not approved for standalone evaluation of the following events and must have separate EM approvals: peak containment pressure (such as following an IORV), control rod ejection accidents, non-LOCA events (other than IORV), return to power analysis assuming the worst-case stuck control rod, and evaluation of the long-term cooling phase.

6. Single Failures, Electrical Power Assumptions (ac/dc) and Need for Operator Actions Not Approved Within this Methodology.

An applicant or licensee seeking to apply this methodology to a design must receive a separate approval through that design review for the single failures, electrical power assumptions (ac/dc) or the need for operator actions necessary to mitigate design basis events necessary to consider for the evaluation of LOCA events, or the IORV events.

7. [[ ]].

[[

]]. Sections 5.5.2 and 5.5.3.2 of this SER describes the basis for this limitation.

8. High Flow CHF Correlation Range.

Application of the high-flow CHF correlation is limited to its range of applicability as identified in Table 5.2-1. Sections 5.5.5.2.2.2 and 5.5.5.1 of this SER describe the basis for this limitation.

Table 5.2-1. High-Flow CHF Correlation Range of Applicability

Parameter Min Max

Pressure [[ ]] [[ ]]

Inlet Subcooling [[ ]] [[ ]]

Mass Flux [[ ]] [[ ]]

9. CHFR Minimum Value.

The high-flow CHFR limit of 1.13 is required for analyses of IORV events for high-flow conditions. The low-flow CHFR limit of 1.37 is required for analyses of IORV events for low-flow conditions. The high-flow and low-flow CHFR limit of 1.29 is required for analyses of LOCA events for high-flow and low-flow conditions. Sections 5.5.4.3 and 5.5.2.2.2 of this SER describe the basis for this condition.

7.0 CONCLUSION

This SER documents the results of the technical evaluation of TR-0516-49422-P, “Loss-of- Coolant Accident Evaluation Model,” Revision 2. The NRC staff finds that the proposed methodology is acceptable for meeting the requirements of 10 CFR 50.46 and Appendix K evaluated in this TR, for evaluation of the ECCS performance in the NuScale NPM for design basis LOCAs, subject to the limitations, conditions, and restrictions identified in Section 6.0 above. The NRC staff finds the NuScale LOCA EM appropriate for determining CHF and CLL results, excluding peak cladding temperature, clad oxidation and core wide clad oxidation, but requires that this information, along with the worst break minimum liquid level in the vessel above the top of the active fuel be reported on a plant specific application, which uses this version of the NuScale LOCA EM. The NRELAP5 computer code is also determined to be applicable to predict peak containment pressure and temperature when referenced in a separately approved EM, subject to specific modeling requirements necessary for prediction of these FOMs. The NRC staff finds the CHF modeling described in TR-0516-49422, Revision 2, acceptable subject to the limitations and conditions identified in Section 6.0 of this SER. Therefore, the NRELAP5 computer code and the NPM model are determined to be acceptable to evaluate the MCHFR for IORV and LOCA events.

For uses other than that intended and approved as part of the NuScale LOCA methodology, the process and all of its elements, including a description of the intended use and justification, must be submitted to the NRC for review and approval. The NRC staff emphasizes that the criterion for acceptable ECCS performance following all LOCA break sizes is that the CLL in the RPV remain above the top of the active fuel and the CHF limit be met. TR-0516-49422-P, Revision 2, constitutes a separate and unique methodology, and as such, any other version 92 derived from this TR, such as an update designated by a new revision number, amendment number, addendum number or equivalent designation would constitute a definition of a new methodology requiring the NRC staff’s review and acceptance prior to a generic application and prior to any specific plant licensing application of a new methodology derived from this TR.

93 Loss-of-Coolant Accident Evaluation Model

TR-0516-49422-NP-A Rev 2

Section B

TR-0516-49422-NP-A Rev. 2

Loss-of-Coolant Accident Evaluation Model

July 2020 Revision 2 Docket: PROJ0769

TR-0516-49422-NP-A Rev. 2

This report has been prepared by NuScale Power, LLC and bears a NuScale Power, LLC, copyright notice. No right to disclose, use, or copy any of the information in this report, other than by the U.S. Nuclear Regulatory Commission (NRC), is authorized without the express, written permission of NuScale Power, LLC.

The NRC is permitted to make the number of copies of the information contained in this report that is necessary for its internal use in connection with generic and plant-specific reviews and approvals, as well as the issuance, denial, amendment, transfer, renewal, modification, suspension, revocation, or violation of a license, permit, order, or regulation subject to the requirements of 10 CFR 2.390 regarding restrictions on public disclosure to the extent such information has been identified as proprietary by NuScale Power, LLC, copyright protection notwithstanding. Regarding nonproprietary versions of these reports, the NRC is permitted to make the number of copies necessary for public viewing in appropriate docket files in public document rooms in Washington, DC, and elsewhere as may be required by NRC regulations. Copies made by the NRC must include this copyright notice and contain the proprietary marking if the original was identified as proprietary.

TR-0516-49422-NP-A Rev. 2

Department of Energy Acknowledgement and Disclaimer

This material is based upon work supported by the Department of Energy under Award Number DE-NE0008928.

This report was prepared as an account of work sponsored by an agency of the United States Government. Neither the United States Government nor any agency thereof, nor any of their employees, makes any warranty, express or implied, or assumes any legal liability or responsibility for the accuracy, completeness, or usefulness of any information, apparatus, product, or process disclosed, or represents that its use would not infringe privately owned rights. Reference herein to any specific commercial product, process, or service by trade name, trademark, manufacturer, or otherwise does not necessarily constitute or imply its endorsement, recommendation, or favoring by the United States Government or any agency thereof. The views and opinions of authors expressed herein do not necessarily state or reflect those of the United States Government or any agency thereof.

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CONTENTS Abstract ...... 1 Executive Summary ...... 3 1.0 Introduction ...... 6 1.1 Purpose ...... 6 1.2 Scope ...... 6 1.3 Abbreviations and Definitions ...... 7 2.0 Background ...... 11 2.1 Loss-of-Coolant Accident Evaluation Model Roadmap ...... 11 2.2 Regulatory Requirements ...... 15 2.2.1 10 CFR 50.46 Loss-of-Coolant Accident Acceptance Criteria ...... 15 2.2.2 NuScale Loss-of-Coolant Accident Evaluation Model Acceptance Criteria ...... 15 2.2.3 10 CFR 50 Appendix K ...... 16 2.2.4 Other Requirements ...... 34 3.0 NuScale Power Module Description and Operations ...... 35 3.1 General Plant Design ...... 35 3.2 Plant Operation ...... 38 3.3 Safety-Related System Operation ...... 39 3.3.1 Emergency Core Cooling System ...... 40 3.3.2 Decay Heat Removal System ...... 41 4.0 Phenomena Identification and Ranking ...... 42 4.1 Phenomena Identification and Ranking Process ...... 42 4.2 Loss-of-Coolant Accident Scenarios ...... 43 4.3 Figures of Merit ...... 44 4.4 Definitions of Importance and Knowledge Level Rankings ...... 45 4.5 Systems, Structures, and Components ...... 45 4.6 High-Ranked Phenomena ...... 46 4.6.1 Discussion of Phenomena Ranked High Importance ...... 48 4.7 Phenomena Identification and Ranking Table Summary ...... 51 5.0 Evaluation Model Description ...... 53 5.1 NRELAP5 Loss-of-Coolant Accident Model for the NuScale Power Module ...... 53 5.1.1 General Model Nodalization ...... 53

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5.1.2 Reactor Coolant System ...... 56 5.1.3 Helical Coil Steam Generators ...... 61 5.1.4 Containment Vessel and Reactor Pool ...... 61 5.1.5 Chemical and Volume Control System ...... 62 5.1.6 Secondary System ...... 63 5.1.7 Decay Heat Removal System ...... 63 5.1.8 NRELAP5 Modeling Options ...... 63 5.1.9 Time Step Size Control ...... 66 5.2 Analysis Setpoints and Trips ...... 67 5.3 Initial Plant Conditions ...... 72 5.4 Loss-of-Coolant Accident Break Spectrum ...... 72 5.4.1 Break Location ...... 73 5.4.2 Break Configuration and Size ...... 73 5.4.3 Single Failures ...... 74 5.4.4 Loss of Power ...... 75 5.4.5 Decay Heat Removal System Availability ...... 76 5.5 Sensitivity Studies ...... 76 6.0 NRELAP5 Code Description ...... 77 6.1 Quality Assurance Requirements ...... 78 6.2 NRELAP5 Hydrodynamic Model ...... 79 6.2.1 Field Equations ...... 79 6.2.2 State Relations ...... 82 6.2.3 Flow Regime Maps ...... 83 6.2.4 Momentum Closure Relations ...... 85 6.2.5 Heat Transfer ...... 91 6.3 Heat Structure Models ...... 93 6.4 Point Reactor Kinetics Model ...... 96 6.5 Trips and Control System Models ...... 97 6.6 Special Solution Techniques ...... 98 6.6.1 Choked Flow ...... 98 6.6.2 Abrupt Area Change ...... 101 6.6.3 Counter Current Flow Limitation ...... 102 6.7 Helical Coil Steam Generator Component ...... 104

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6.7.1 Helical Coil Tube Friction ...... 105 6.7.2 Helical Coil Tube Heat Transfer ...... 106 6.8 Wall Heat Transfer and Condensation ...... 108 6.8.1 NRELAP5 Solution Approach for Wall Condensation Heat Transfer ..... 109 6.8.2 Wall Condensation Correlation ...... 113 6.9 Interfacial Drag in Large Diameter Pipes...... 115 6.10 Fission Decay Heat and Actinide Models ...... 116 6.11 Critical Heat Flux Models ...... 118 6.11.1 {{ }}2(a),(c) ...... 119 6.11.2 Implementation of Critical Heat Flux correlations ...... 121 6.11.3 {{ }}2(a),(c) ...... 121 6.11.4 {{ }}2(a),(c) ...... 125 7.0 NRELAP5 Assessments ...... 127 7.1 Assessment Methodology ...... 127 7.2 Legacy Test Data ...... 129 7.2.1 Ferrell-McGee ...... 129 7.2.2 GE Level Swell (1 ft) ...... 133 7.2.3 GE Level Swell (4 ft) ...... 141 7.2.4 KAIST ...... 146 7.2.5 FRIGG ...... 153 7.2.6 FLECHT-SEASET ...... 160 7.2.7 SemiScale (S-NC-02 and S-NC-10) ...... 169 7.2.8 Wilson Bubble Rise ...... 174 7.2.9 Marviken Jet Impingement Test (JIT) 11 ...... 179 7.2.10 Bankoff Perforated Plate ...... 182 7.2.11 Marviken Critical Flow Test 22 and 24 ...... 184 7.3 NuScale Stern Critical Heat Flux Tests ...... 192 7.3.1 Facility Description ...... 193 7.3.2 Experimental Procedure ...... 194 7.3.3 Phenomenon Addressed ...... 195 7.3.4 Parameter Ranges Assessed ...... 195 7.3.5 Special Analysis Techniques ...... 196 7.3.6 Assessment Results ...... 196

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7.4 NuScale SIET Steam Generator Tests ...... 197 7.4.1 SIET Tests ...... 197 7.4.2 SIET Fluid-Heated Test ...... 207 7.5 NuScale NIST-1 Test Assessment Cases ...... 217 7.5.1 Test Facility Description ...... 218 7.5.2 Facility NRELAP5 Model ...... 224 7.5.3 Facility Test Matrix ...... 224 7.5.4 Separate Effect High Pressure Condensation Tests ...... 226 7.5.5 Natural Circulation Test at Power ...... 240 7.5.6 Chemical and Volume Control System Loss-of-Coolant Accident Integral Effects Tests ...... 244 7.5.7 Pressurizer Spray Supply Line Loss-of-Coolant Accident Integral Effects Test ...... 263 7.5.8 Spurious Reactor Vent Valve Opening Test ...... 271 8.0 Assessment of Evaluation Model Adequacy ...... 281 8.1 Adequacy Demonstration Overview ...... 281 8.2 Evaluation of Models and Correlations (Bottom-Up Assessment) ...... 281 8.2.1 Important Models and Correlations ...... 283 8.2.2 {{ }}2(a),(c) ...... 292 8.2.3 {{ }}2(a),(c) ...... 297 8.2.4 {{ }}2(a),(c) ...... 299 8.2.5 {{ }}2(a),(c) ...... 302 8.2.6 {{ }}2(a),(c) ...... 303 8.2.7 {{ }}2(a),(c) ...... 308 8.2.8 {{ }}2(a),(c) ...... 309 8.2.9 {{ }}2(a),(c) ...... 315 8.2.10 Flashing ...... 316 8.2.11 {{ }}2(a),(c) ...... 317 8.2.12 {{ }}2(a),(c) ...... 321 8.2.13 {{ }}2(a),(c) ...... 322 8.2.14 {{ }}2(a),(c) ...... 323

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8.2.15 {{ }}2(a),(c) ...... 326 8.2.16 {{ }}2(a),(c) ...... 326 8.2.17 {{ }}2(a),(c) ...... 327 8.2.18 {{ }}2(a),(c) ...... 330 8.2.19 {{ }}2(a),(c) ...... 332 8.2.20 {{ }}2(a),(c) ...... 335 8.2.21 {{ }}2(a),(c) ...... 336 8.2.22 {{ }}2(a),(c) ...... 338 8.3 Evaluation of Integral Performance (Top-Down Assessment) ...... 339 8.3.1 Review of Code Governing Equations and Numerics ...... 340 8.3.2 NuScale Facility Scaling ...... 342 8.3.3 Assessment of NuScale Facility Integral Effects Test Data ...... 359 8.3.4 Evaluation of NuScale Integral Effects Tests Distortions and NRELAP5 Scalability ...... 360 8.3.5 Calculation of Peak CNV pressure ...... 369 8.4 Summary of Adequacy Findings ...... 370 8.4.1 Findings from Bottom-Up Evaluation ...... 370 8.4.2 Findings from Top-Down Evaluation ...... 379 8.4.3 Summary of Biases and Uncertainties ...... 384 9.0 Loss-of-Coolant Accident Calculations ...... 385 9.1 Loss-of-Coolant Accident Progression in the NuScale Power Module ...... 385 9.1.1 Liquid Space Break...... 385 9.1.2 Steam Space Break ...... 393 9.2 Break Size ...... 397 9.3 Decay Heat Removal System Availability ...... 401 9.4 Power Availability ...... 403 9.5 Single Failure ...... 403 9.6 Sensitivity Studies ...... 405 9.6.1 Model Nodalization ...... 405 9.6.2 Time-Step Size Selection ...... 409 9.6.3 Counter Current Flow Limitation Behavior on Pressurizer Baffle Plate ...... 411 9.6.4 Emergency Core Cooling System Valve Parameters ...... 412

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9.6.5 Initial Reactor Pool Temperature ...... 414 9.6.6 Core Power Distribution ...... 416 9.7 Loss-of-Coolant Accident Calculation Summary ...... 419 10.0 Conclusions ...... 421 11.0 References ...... 424 Appendix A. Input for NuScale Power Module Loss-of-Coolant Accident Model ...... 430 Appendix B. Evaluation Model for Inadvertent Opening of RPV Valves ...... 437 Appendix C. Spurious Reactor Recirculation Valve Opening Integral Effects Test ...... 509

TABLES Table 1-1. Abbreviations ...... 7 Table 1-2. Definitions ...... 9 Table 2-1. Evaluation model development and assessment process steps and the associated sections in this document ...... 13 Table 2-2. 10 CFR 50 Appendix K required and acceptable features compliance ...... 18 Table 4-1. Importance rankings ...... 45 Table 4-2. Knowledge levels ...... 45 Table 4-3. Systems, structures, and components ...... 45 Table 4-4. High-ranked phenomena ...... 46 Table 5-1. Default junction options for the NRELAP5 loss-of-coolant accident model ...... 64 Table 5-2. Default volume options for the NRELAP5 loss-of-coolant accident model ...... 65 Table 5-3. NuScale Power Module safety-related system measurement parameters ...... 68 Table 5-4. Safety-related system actuation signals ...... 69 Table 5-5. Safety-related analysis signal delays ...... 70 Table 5-6. Plant initial conditions ...... 72 Table 5-7. Summary of analyzed break sizes...... 74 Table 5-8. NuScale Power Module valve fail-safe positions with loss of DC power ...... 75 Table 6-1. Extended Shah dimensionless vapor velocity transition criteria ...... 115 Table 6-2. Extended Shah condensation heat transfer coefficients dependent on regime ...... 115 Table 6-3. ANS 1973 11-group fission decay constants ...... 116 Table 6-4. ANS-79 actinide model constants...... 118 Table 6-5. Coefficient of revised pressure correction term in Equation 6-108 ...... 123 Table 6-6. {{ }}2(a),(c) critical heat flux correlation application range ...... 124 Table 7-1. NRELAP5 loss-of-coolant accident assessment matrix ...... 128 Table 7-2. Summary of Ferrell-McGee experimental test data range ...... 131 Table 7-3. Summary of GE 1 ft. vessel level swell experiments ...... 135 Table 7-4. Range of KAIST test data ...... 149 Table 7-5. Range of Stern steady state critical heat flux data ...... 195 Table 7-6. Facility high priority tests for NRELAP5 code validation ...... 225 Table 8-1. Dominant NRELAP5 models and correlations...... 284 Table 8-2. NuScale Power Module range of process parameters ...... 287 Table 8-3. Range of NuScale Power Module geometric parameters ...... 289 Table 8-4. Marviken range of parameters compared to the NuScale Power Module ...... 295

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Table 8-5. Ferrell-McGee range of parameters compared to the NuScale Power Module ...... 298 Table 8-6. Dimensions of NuScale Power Module, NIST-1 and Bankoff pressurizer plate ...... 301 Table 8-7. Range of riser interphase friction - separate effects tests and NuScale Power Module ...... 307 Table 8-8. Ranges of key parameters for core interphase friction - separate effects tests and plant ...... 329 Table 8-9. Range of key parameters for core flow – separate effects tests and plant ...... 331 Table 8-10. Range of key parameters for core boiling - separate effects tests and plant .... 334 Table 8-11. Range of key parameters for subcooling boiling and separate effects tests and plant ...... 338 Table 8-12. Scaling factors for NIST-1 facility ...... 348 Table 8-13. Mass Flow Paths for NPM and NIST-1 (RCS and CNV) ...... 355 Table 8-14. Heat Flow Paths for RCS in NPM and NIST-1 ...... 355 Table 8-15. Heat Flow Paths for Containment in NPM and NIST-1 ...... 355 Table 8-16. Description of π Groups for the RCS Mass/Energy Balance ...... 356 Table 8-17. Description of π Groups for the Containment Mass/Energy Balance ...... 356 Table 8-18. Summary of bottom-up evaluation of NRELAP5 models and correlations ...... 371 Table 8-19. Applicability summary for high-ranked phenomena ...... 380 Table 9-1. Event table for 100 percent reactor coolant system injection line break ...... 387 Table 9-2. Event table for {{ }}2(a),(c) ...... 394 Table 9-3. Number of volumes in reactor pressure vessel and containment vessel nodalization ...... 406 Table A-1. Core input parameters ...... 431 Table A-2. Initial conditions for loss-of-coolant accident analysis ...... 433 Table A-3. Safety signal actuation setpoints and delays...... 434 Table A-4. Break spectrum parameters ...... 436 Table B-1. Anticipated Operational Occurrence Regulatory Acceptance Criteria ...... 442 Table B-2. KATHY Inlet Boundary Condition Ranges ...... 451 Table B-3. Changes to LOCA EM for IORV EM ...... 455 Table B-4. Comparison of LOCA and IORV initial conditions ...... 459 Table B-5. NIST-1 Spurious RVV Test Differences ...... 462 Table B-6. Sequence of Events for RVV opening without loss of normal AC or DC power ...... 470 Table B-7. Sequence of Events for RRV opening with loss of normal AC and DC power ...... 473 Table B-8. IORV analysis initial conditions ...... 475 Table B-9. Results for RVV cases with power available ...... 477 Table B-10. Results for RRV cases with power available ...... 478 Table B-11. Results for RSV cases with power available ...... 479 Table B-12. Gap conductance results ...... 480 Table B-13. Axial power shape results ...... 480 Table B-14. ECCS valve capacity results ...... 481 Table B-15. ECCS valve stroke time results ...... 482 Table B-16. DHRS operation results ...... 482 Table B-17. Single failure results ...... 483 Table B-18. Electric power availability results ...... 484

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FIGURES Figure 2-1. Evaluation model development and assessment process ...... 12 Figure 3-1. A single NuScale Power Module during normal operation ...... 36 Figure 3-2. Schematic of NuScale Power Module decay heat removal system and emergency core cooling system during operation ...... 38 Figure 5-1. Noding diagram of NRELAP5 loss-of-coolant accident input model for NuScale Power Module ...... 55 Figure 6-1. Schematic of vertical flow-regime map indicating transitions ...... 84 Figure 6-2. NRELAP5 boiling and condensing curves ...... 93 Figure 6-3. {{ }}2(a),(c) ...... 109 Figure 6-4. NRELAP5 ANS 1973 implemented fission decay heat curve ...... 117 Figure 6-5. NRELAP5 ANS-79 implemented actinide heat curve ...... 118 Figure 6-6. {{ }}2(a),(c) ...... 120 Figure 6-7. {{ }}2(a),(c) ...... 124 Figure 7-1. Schematic of the Ferrell-McGee test section ...... 130 Figure 7-2. Predicted versus measured pressure drop for selected contraction tests ...... 133 Figure 7-3. Schematic of the GE 1 ft. blowdown vessel ...... 134 Figure 7-4. GE level swell 1 ft. vessel pressure versus time ...... 137 Figure 7-5. GE level swell 1 ft. vessel void fraction versus elevation at 10 seconds ...... 138 Figure 7-6. GE level swell 1 ft. vessel void fraction versus elevation at 40 seconds ...... 139 Figure 7-7. GE level swell 1 ft. vessel void fraction versus elevation at 100 seconds ...... 140 Figure 7-8. GE level swell 1 ft. vessel void fraction versus elevation at 160 seconds ...... 141 Figure 7-9. Schematic of the GE 4 ft. blowdown vessel ...... 142 Figure 7-10. GE level swell 4-ft vessel pressure versus time ...... 144 Figure 7-11. GE level swell 4-ft vessel void fraction versus elevation at 5 seconds ...... 144 Figure 7-12. GE level swell 4-ft vessel void fraction versus elevation at 10 seconds ...... 145 Figure 7-13. GE level swell 4-ft vessel void fraction versus elevation at 20 seconds ...... 145 Figure 7-14. Schematic of KAIST test facility ...... 147 Figure 7-15. Schematic of the KAIST test section ...... 148 Figure 7-16. Measured versus predicted heat transfer coefficient ...... 150 Figure 7-17. KAIST and NRELAP5 axial heat transfer coefficient ...... 151 Figure 7-18. KAIST and NRELAP5 axial inner wall temperature ...... 152 Figure 7-19. KAIST and NRELAP5 axial liquid mass flow rate ...... 153 Figure 7-20. FRIGG-4 experimental loop ...... 155 Figure 7-21. FRIGG-4 36 rod test section ...... 156 Figure 7-22. FRIGG-4 zones for evaluation of radial void distribution ...... 157 Figure 7-23. FRIGG mean void data of NRELAP5 versus Test 613123 data ...... 158 Figure 7-24. FRIGG mean void data of NRELAP5 versus Test 613130 data ...... 159 Figure 7-25. FRIGG mean void data of NRELAP5 versus Test 613010 data ...... 159 Figure 7-26. FRIGG mean void data of NRELAP5 versus Test 613118 data ...... 160 Figure 7-27. FLECHT-SEASET experimental facility ...... 161 Figure 7-28. FLECHT-SEASET level 1 void fraction versus time – Test 35557...... 162 Figure 7-29. FLECHT-SEASET level 2 void fraction versus time – Test 35557...... 163 Figure 7-30. FLECHT-SEASET level 3 void fraction versus time – Test 35557...... 164 Figure 7-31. FLECHT-SEASET level 4 void fraction versus time – Test 35557...... 165 Figure 7-32. FLECHT-SEASET level 1 collapsed water level versus time – Test 35557...... 166 Figure 7-33. FLECHT-SEASET level 2 collapsed water level versus time – Test 35557...... 167 Figure 7-34. FLECHT-SEASET level 3 collapsed water level versus time – Test 35557...... 168

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Figure 7-35. FLECHT-SEASET level 4 collapsed water level versus time – Test 35557...... 169 Figure 7-36. Semiscale Mod-2A single (intact) loop test facility configuration ...... 170 Figure 7-37. S-NC-2 30 kW average mass flow rate versus percent inventory ...... 172 Figure 7-38. S-NC-2 60 kW average mass flow rate versus percent inventory ...... 173 Figure 7-39. S-NC-10 100 kW average mass flow rate versus percent inventory ...... 174 Figure 7-40. Schematic of Wilson bubble rise test facility ...... 175 Figure 7-41. NRELAP5 and Wilson void fraction versus superficial velocity at 600 psig (4.14 MPa) ...... 176 Figure 7-42. NRELAP5 and Wilson void fraction versus superficial velocity 1,000 psig (6.89 MPa) ...... 177 Figure 7-43. NRELAP5 and Wilson void fraction versus superficial velocity 2,000 psig (13.8 MPa) ...... 178 Figure 7-44. Predicted versus measured area averaged void fraction (all cases) ...... 179 Figure 7-45. Marviken jet impingement test facility ...... 180 Figure 7-46. Marviken jet impingement test 11 flowrate ...... 181 Figure 7-47. Marviken jet impingement test 11 density ...... 182 Figure 7-48. Schematic of Bankoff counter current flow apparatus (from Reference 68) ...... 183 Figure 7-49. Superficial vapor velocity versus superficial liquid velocity ...... 184 Figure 7-50. Schematic of the Marviken pressure vessel ...... 185 Figure 7-51. Discharge pipe dimensions and instrument locations ...... 186 Figure 7-52. Measured versus calculated mass flow rate for Marviken critical flow test 22 ...... 189 Figure 7-53. Marviken critical flow test 22 comparison to calculated mixture density ...... 190 Figure 7-54. Measured versus calculated mass flow rate for Marviken critical flow test 24 ...... 191 Figure 7-55. Marviken critical flow test 24 mixture density and calculated mixture density ... 192 Figure 7-56. U1 & C1 (left) versus U2 (right) radial layout ...... 193 Figure 7-57. Stern test section axial layout ...... 194 Figure 7-58. Predicted versus measured Stern power ...... 197 Figure 7-59. SIET electrically-heated test instrumentation diagram ...... 199 Figure 7-60. Time averaged wall temperature profile for coil 2 test TD0015 ...... 201 Figure 7-61. Time averaged wall temperature profile for coil 2 test TD0003 ...... 202 Figure 7-62. SIET electrically-heated test differential pressure for all coil 1 diabatic tests .... 203 Figure 7-63. SIET electrically-heated test differential pressure for all coil 2 diabatic tests .... 204 Figure 7-64. SIET electrically-heated test differential pressure for all coil 3 diabatic tests .... 205 Figure 7-65. SIET electrically-heated test fluid temperatures for all coil 1 diabatic tests ...... 206 Figure 7-66. SIET electrically-heated test wall temperature for all coil 1 diabatic tests ...... 207 Figure 7-67. SIET fluid-heated test adiabatic primary differential pressure ...... 210 Figure 7-68. SIET fluid-heated test diabatic test primary differential pressure ...... 211 Figure 7-69. SIET fluid-heated test diabatic test primary temperature ...... 212 Figure 7-70. Comparison of wall temperatures in TD0001 (Case 1A) ...... 213 Figure 7-71. Comparison of wall temperatures in TD0005 (Case 1A) ...... 214 Figure 7-72. Comparison of wall temperatures in TD0015 (Case 1A) ...... 215 Figure 7-73. Comparison of primary and secondary side fluid temperatures in TD0001 (Case 1A) ...... 216 Figure 7-74. Comparison of primary and secondary side fluid temperatures in TD0005 (Case 1A) ...... 217 Figure 7-75. Schematic of NuScale integral test facility and NRELAP5 nodalization ...... 219 Figure 7-76. HP-02 Run 1 containment vessel pressure response ...... 229

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Figure 7-77. HP-02 Run 1 containment vessel collapsed level response ...... 230 Figure 7-78. HP-02 Run 1 upper containment vessel fluid temperature response (in vapor space) ...... 231 Figure 7-79. HP-02 Run 1 upper cooling pool vessel temperature response ...... 232 Figure 7-80. HP-02 Run 2 containment vessel pressure response ...... 233 Figure 7-81. HP-02 Run 2 containment vessel collapsed level response ...... 234 Figure 7-82. HP-02 Run 2 upper containment vessel fluid temperature response (in vapor space) ...... 235 Figure 7-83. HP-02 Run 2 upper cooling pool temperature response ...... 236 Figure 7-84. HP-02 Run 3 containment vessel pressure response ...... 237 Figure 7-85. HP-02 Run 3 containment vessel collapsed level response ...... 238 Figure 7-86. HP-02 Run 3 upper containment vessel fluid temperature response (in vapor space) ...... 239 Figure 7-87. HP-02 Run 3 upper cooling pool temperature response ...... 240 Figure 7-88. HP-05 NIST-1 averaged mass flowrate and NRELAP5 results ...... 242 Figure 7-89. HP-05 NIST-1 averaged core inlet temperature and NRELAP5 results ...... 243 Figure 7-90. HP-05 NIST-1 averaged core outlet temperature and NRELAP5 results ...... 244 Figure 7-91. NIST-1 HP-06 NRELAP5 chemical and volume control system discharge line break mass flow rate ...... 247 Figure 7-92. NIST-1 HP-06 break orifice differential pressure ...... 248 Figure 7-93. NIST-1 HP-06 primary mass flow rate ...... 249 Figure 7-94. NIST-1 HP-06 pressurizer level comparison ...... 250 Figure 7-95. NIST-1 HP-06 reactor pressure vessel level comparison ...... 251 Figure 7-96. NIST-1 HP-06 containment vessel level comparison ...... 252 Figure 7-97. NIST-1 HP-06 containment vessel pressure comparison ...... 253 Figure 7-98. NIST-1 HP-06 containment vessel pressure comparison ...... 254 Figure 7-99. NIST-1 HP-06 primary pressure comparison ...... 254 Figure 7-100. Comparison of core power in HP-06 and HP-06b tests with the NuScale Power Module decay power after reactor trip (scaled) ...... 255 Figure 7-101. NIST-1 HP-06b primary pressure comparison ...... 256 Figure 7-102. NIST-1 HP-06b containment vessel pressure comparison ...... 257 Figure 7-103. NIST-1 HP-06b reactor pressure vessel level comparison ...... 258 Figure 7-104. NIST-1 HP-06b containment vessel level comparison ...... 259 Figure 7-105. Comparison of NIST-1 HP-06 and HP-06b reactor pressure vessel pressure ...... 260 Figure 7-106. Comparison of NIST-1 HP-06 and HP-06b containment vessel pressure ...... 261 Figure 7-107. Comparison of NIST-1 HP-06 and HP-06b reactor pressure vessel level ...... 262 Figure 7-108. Comparison of NIST-1 HP-06 and HP-06b containment vessel level ...... 263 Figure 7-109. Comparison of core power in HP-07 with the NuScale Power Module power (fission and decay) after reactor trip (scaled) ...... 265 Figure 7-110. NIST-1 HP-07 pressurizer spray supply line break discharge mass flow rate ... 266 Figure 7-111. NIST-1 HP-07 primary mass flow rate ...... 267 Figure 7-112. NIST-1 HP-07 reactor pressure vessel level response comparison with data ...... 268 Figure 7-113. NIST-1 HP-07 containment vessel level response ...... 269 Figure 7-114. NIST-1 HP-07 containment vessel pressure comparison ...... 270 Figure 7-115. NIST-1 HP-07 primary pressure comparison ...... 271 Figure 7-116. Comparison of HP-09 core power with the scaled NuScale Power Module fission and decay power ...... 273

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Figure 7-117. NIST-1 HP-09 valve mass flow rate ...... 274 Figure 7-118. NIST-1 HP-09 pressurizer pressure comparison ...... 275 Figure 7-119. NIST-1 HP-09 pressurizer pressure comparison ...... 275 Figure 7-120. NIST-1 HP-09 containment vessel pressure comparison ...... 276 Figure 7-121. NIST-1 HP-09 containment vessel pressure comparison ...... 277 Figure 7-122. NIST-1 HP-09 pressurizer level comparison ...... 278 Figure 7-123. NIST-1 HP-09 reactor pressure vessel level comparison ...... 279 Figure 7-124. NIST-1 HP-09 reactor pressure vessel level comparison ...... 280 Figure 8-1. CNV wall heat transfer modes ...... 311 Figure 8-2. Thermal resistance network between CNV and UHS ...... 311 Figure 8-3. Transient void fraction in node 5 for the GE 4-ft level swell test ...... 319 Figure 8-4. Transient void fraction in node 4 for the GE 4-ft level swell test ...... 319 Figure 8-5. Transient void fraction in node 6 for the GE 1-ft level swell test ...... 320 Figure 8-6. Transient void fraction in node 5 for the GE 1-ft level swell test ...... 320 Figure 8-7. General design framework for the NuScale Integral System Test facility ...... 343 Figure 8-8. Flow diagram for the hierarchical, two-tiered scaling analysis (NUREG/CR-5809) ...... 345 Figure 8-9. NuScale system breakdown into hierarchical levels and primary operational modes to be scaled ...... 346 Figure 8-10. Scaling analysis flow diagram for single-phase primary loop natural circulation ...... 349 Figure 8-11. Scaling analysis flow diagram for reactor coolant system depressurization ...... 350 Figure 8-12. Scaling analysis flow diagram for containment pressurization ...... 351 Figure 8-13. Scaling analysis flow diagram for long-term recirculation cooling mode ...... 352 Figure 8-14. Scaling analysis flow diagram for Reactor Building pool heat-up ...... 353 Figure 8-15. Comparison of HP-05 feedwater flow to test data ...... 363 Figure 8-16. Comparison of HP-05 reactor pressure vessel flow to test data ...... 363 Figure 8-17. Comparison of HP-05 upper riser inlet temperature to test data ...... 365 Figure 8-18. Comparison of HP-05 core inlet temperature to test data ...... 365 Figure 9-1. Break and emergency core cooling system valve flows to the containment vessel for 100 percent injection line break ...... 388 Figure 9-2. Integrated energy for 100 percent injection line break ...... 388 Figure 9-3. Reactor pressure vessel and containment vessel pressure for 100 percent injection line break ...... 389 Figure 9-4. Comparison of collapsed liquid levels in reactor pressure vessel, containment vessel, and pressurizer for 100 percent injection line break ...... 390 Figure 9-5. Core flow for 100 percent injection line break ...... 391 Figure 9-6. Core minimum critical heat flux ratio during 100 percent injection line break .... 392 Figure 9-7. Peak cladding and fuel centerline temperature during 100 percent injection line break ...... 392 Figure 9-8. Comparison of pressure for injection line and high point vent line breaks ...... 395 Figure 9-9. Break and emergency core cooling system valve flow during 100 percent high point vent line break ...... 396 Figure 9-10. Collapsed liquid levels during 100 percent high point vent line break ...... 396 Figure 9-11. Minimum critical heat flux ratio during 100 percent high point vent line break ...... 397 Figure 9-12. Peak containment vessel pressure and collapsed level above top of active fuel for different reactor coolant system injection line break sizes ...... 398

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Figure 9-13. Peak containment vessel pressure and collapsed level above top of active fuel for different reactor coolant system discharge line break sizes ...... 399 Figure 9-14. Peak containment vessel pressure and collapsed level above top of active fuel for different high point vent line break sizes ...... 399 Figure 9-15. Peak containment vessel pressure and minimum collapsed liquid level as a function of break location and size ...... 400 Figure 9-16. Minimum critical heat flux ratio for injection line (left) and high point vent line (right) breaks ...... 401 Figure 9-17. Reactor coolant system and containment pressures for reactor coolant system injection line break without decay heat removal system (left) and with decay heat removal system (right) ...... 402 Figure 9-18. Collapsed liquid level for reactor coolant system injection line break without decay heat removal system (left) and with decay heat removal system (right) ...... 402 Figure 9-19. Effect of power availability on peak containment vessel pressure for injection line (left) and high point vent (right) line breaks ...... 403 Figure 9-20. The effect of single failure on peak containment vessel pressure for reactor coolant system injection line (left) and high point vent (right) line breaks ...... 404 Figure 9-21. The effect of single failure on minimum collapsed level for reactor coolant system injection line (left) and high point vent line (right) breaks ...... 405 Figure 9-22. Reactor pressure vessel and containment vessel pressure (left) and collapsed level above top of active fuel (right) for 100 percent reactor coolant system injection line break ...... 407 Figure 9-23. Reactor pressure vessel and containment vessel pressure (left) and collapsed level above top of active fuel (right) for 10 percent reactor coolant system injection line break ...... 407 Figure 9-24. Reactor pressure vessel and containment vessel pressure (left) and collapsed level above top of active fuel (right) for 100 percent high point vent line break ...... 408 Figure 9-25. Hot channel core flow (left) and core critical heat flux ratio (right) during 100 percent reactor coolant system injection line break ...... 409 Figure 9-26. Time-step size sensitivity on reactor and containment vessel pressures and reactor pressure vessel collapsed liquid level for 100 percent reactor coolant system injection line break...... 410 Figure 9-27. Time-step size sensitivity on hot assembly flow and minimum critical heat flux ratio for 100 percent reactor coolant system injection line break ...... 410 Figure 9-28. Time-step size sensitivity on reactor and containment vessel pressures and reactor pressure vessel collapsed liquid level for 100 percent high point vent line break ...... 411 Figure 9-29. Time-step size sensitivity on hot assembly flow and minimum critical heat flux ratio for 100 percent high point vent break ...... 411 Figure 9-30. Effect of counter current flow limitation line slope on levels for 100 percent high point vent line break ...... 412 Figure 9-31. Effect of inadvertent actuation block release pressure on peak containment vessel pressure and minimum collapsed liquid level above top of active fuel as a function of break size for reactor coolant system injection line break ...... 413 Figure 9-32. Effect of reactor recirculation valve size on peak containment vessel pressure and minimum collapsed liquid level for reactor coolant system injection line break ...... 414

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Figure 9-33. Effect of reactor vent valve size on peak containment vessel pressure and minimum collapsed liquid level for reactor coolant system injection line break ...... 414 Figure 9-34. Effect of initial reactor pool temperature on peak containment vessel pressure and minimum collapsed liquid level above top of active fuel for reactor coolant system injection line break...... 415 Figure 9-35. Containment vessel to pool energy transfer at different initial pool temperatures for 100 percent (left) and 10 percent (right) reactor coolant system injection line break ...... 416 Figure 9-36. Generic normalized axial power shapes...... 417 Figure 9-37. Effect of axial power shape on reactor pressure vessel and containment pressures and collapsed liquid level above top of active fuel for reactor coolant system injection line break ...... 418 Figure 9-38. Effect of axial power shape on peak containment vessel pressure and minimum collapsed liquid level above top of active fuel for reactor coolant system injection line break ...... 418 Figure 9-39. Effect of axial power shape on hot assembly flow and minimum critical heat flux ratio during reactor coolant system injection line break ...... 419 Figure B-1. KATHY Test Loop ...... 448 Figure B-2. KATHY NuFuel HTP2TM Test Section Axial Layout ...... 449 Figure B-3. KATHY {{ }}2(a),(c),ECI Radial Layouts ...... 450 Figure B-4. CHF Statistical Methods Flow Chart ...... 452 Figure B-5. HS171 Correlation: Predicted vs Measured Power ...... 454 Figure B-6. HP-43 transient short-term pressurizer pressure comparison ...... 463 Figure B-7. HP-43 transient short-term pressurizer level code-to-data comparison ...... 464 Figure B-8. HP-43 transient short-term RPV code-to-data level comparison ...... 465 Figure B-9. HP-43 transient short-term CNV pressure code-to-data comparison ...... 466 Figure B-10. HP-43 transient short-term spurious RVV orifice mass flow rate code-to-data comparison ...... 467 Figure B-11. HP-43 transient short-term CNV level code-to-data comparison ...... 468 Figure B-12. Inadvertently opened RVV flow (short term) ...... 485 Figure B-13. Inadvertently opened RVV flow (long term) ...... 485 Figure B-14. RPV and CNV pressure (short term) for inadvertent RVV opening ...... 486 Figure B-15. RPV and CNV pressure (long term) for inadvertent RVV opening ...... 486 Figure B-16. RCS flow (short term) for inadvertent RVV opening ...... 487 Figure B-17. RCS flow (long term) for inadvertent RVV opening ...... 488 Figure B-18. RCS temperatures (short term) for inadvertent RVV opening ...... 489 Figure B-19. RCS temperatures (long term) for inadvertent RVV opening ...... 490 Figure B-20. CHFR (short term) for inadvertent RVV opening ...... 490 Figure B-21. CHFR (long term) for inadvertent RVV opening ...... 491 Figure B-22. Transient MCHFR for inadvertent RVV opening ...... 491 Figure B-23. RPV and CNV level (short term) for inadvertent RVV opening ...... 492 Figure B-24. RPV and CNV level (long term) for inadvertent RVV opening ...... 492 Figure B-25. Reactor Power (short term) for inadvertent RVV opening ...... 493 Figure B-26. Net reactivity (short term) for inadvertent RVV opening ...... 493 Figure B-27. SG-1 pressure for inadvertent RVV opening ...... 494 Figure B-28. ECCS (non-inadvertently opened) RVV flow ...... 494 Figure B-29. ECCS (non-inadvertently opened) RRV flow ...... 495 Figure B-30. Fuel temperature (°F) for inadvertent RVV opening (short term) ...... 495

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Figure B-31. Fuel temperature (°F) for inadvertent RVV opening (long term) ...... 496 Figure B-32. Inadvertently opened RRV flow (short term) ...... 497 Figure B-33. Inadvertently opened RRV flow (long term) ...... 497 Figure B-34. RPV and CNV pressure (short term) for inadvertent RRV opening ...... 498 Figure B-35. RPV and CNV pressure (long term) for inadvertent RRV opening ...... 498 Figure B-36. RCS flow (short term) for inadvertent RRV opening ...... 499 Figure B-37. RCS flow (long term) for inadvertent RRV opening ...... 500 Figure B-38. RCS temperatures (short term) for inadvertent RRV opening ...... 501 Figure B-39. RCS temperatures (long term) for inadvertent RRV opening ...... 501 Figure B-40. CHFR (short term) for inadvertent RRV opening ...... 502 Figure B-41. CHFR (long term) for inadvertent RRV opening ...... 502 Figure B-42. Transient MCHFR for inadvertent RRV opening ...... 503 Figure B-43. RPV and CNV level (short term) for inadvertent RRV opening ...... 503 Figure B-44. RPV and CNV level (long term) for inadvertent RRV opening ...... 504 Figure B-45. Reactor Power (short term) for inadvertent RRV opening ...... 504 Figure B-46. Net reactivity (short term) for inadvertent RRV opening ...... 505 Figure B-47. SG-1 pressure for inadvertent RRV opening ...... 505 Figure B-48. ECCS (non-inadvertently opened) RRV flow ...... 506 Figure B-49. Fuel temperature (°F) for inadvertent RRV opening (short term) ...... 506 Figure B-50. Fuel temperature (°F) for inadvertent RRV opening (long term) ...... 507 Figure C-1. NIST-1 HP-49 spurious orifice differential pressure ...... 511 Figure C-2. NIST-1 HP-49 primary mass flow rate ...... 512 Figure C-3. NIST-1 HP-49 pressurizer level comparison ...... 512 Figure C-4. NIST-1 HP-49 reactor pressure vessel level comparison ...... 513 Figure C-5. NIST-1 HP-49 containment vessel level comparison ...... 513 Figure C-6. NIST-1 HP-49 containment vessel peak pressure comparison ...... 514 Figure C-7. NIST-1 HP-49 containment vessel pressure comparison ...... 514 Figure C-8. NIST-1 HP-49 primary pressure comparison ...... 515

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Abstract

NuScale Power, LLC (NuScale) has developed a small modular reactor (SMR) that supports operation of up to 12 NuScale Power Modules (NPM) at a specific site. Each NPM is an advanced, light-water, integrated pressurized water reactor (PWR) using natural circulation for primary coolant flow. Each NPM has an independent nuclear steam supply system (NSSS), a standard steam power conversion system, and a compact steel containment vessel (CNV). In the NPM design, all primary components are integral to the reactor pressure vessel (RPV), which eliminates most of the reactor piping found on conventional PWRs, thereby reducing the possibility of a pipe rupture that would result in a loss-of-coolant accident (LOCA).

NuScale is requesting Nuclear Regulatory Commission (NRC) review and approval to use the LOCA evaluation model (EM) described in this report for analyses of design-basis LOCA events in the NPM. The NuScale LOCA EM has been developed using the evaluation model development and assessment process (EMDAP) of “Transient and Accident Analysis Methods,” Regulatory Guide (RG) 1.203 (Reference 1), and it adheres to the applicable requirements of “ECCS Evaluation Models,” 10 CFR 50 Appendix K (Reference 2), and “Acceptance Criteria for Emergency Core Cooling Systems for Light-Water Nuclear Power Reactors,” 10 CFR 50.46 (Reference 3). This topical report is not intended to provide final design values or results; rather, example values for the various evaluations are provided for illustrative purposes in order to aid the reader’s understanding of the context of the application of the NuScale LOCA EM.

Validation and verification of the LOCA EM and NRELAP5 code has been performed in accordance with the EMDAP. A phenomena identification and ranking table (PIRT), which identifies the important phenomena and processes occurring in the NPM during a LOCA, was developed by gathering and ranking expert evaluations of phenomena that could occur in the NPM during a LOCA. Twenty-one (21) phenomena were identified as important to capture in the NuScale LOCA EM.

Extensive NRELAP5 code validation was performed to ensure that the LOCA EM is applicable for important phenomena and processes over the range encountered in the NPM LOCA. The validation suite includes many legacy separate effects tests (SETs) and integral effects tests (IETs), as well as many SETs and IETs developed and run specifically for the NPM application.

The EMDAP requires an applicability demonstration of the NRELAP5 code and tests. A unique aspect of the demonstration provided for the NPM is the comparison of NRELAP5 simulations of LOCA to NuScale Integral System Test Facility (NIST-1) test data and NRELAP5 simulation of the same LOCA in the NPM. The reasonable-to-excellent agreement obtained by these comparisons establishes the applicability of NRELAP5 to accurately predict LOCA phenomena at both the NIST-1 and NPM scales.

This topical report provides an example application of the LOCA EM in order to aid the reader’s understanding of the context of the application of the NuScale LOCA EM. These calculations

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are presented for break spectra that cover a range of break locations, break sizes, single failures, equipment unavailability, and initial and boundary conditions. The methodology in this report is also used to support analyses for Non-LOCA events, containment peak pressure analysis, long term cooling evaluation and inadvertent ECCS actuation.

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Executive Summary

NuScale Power, LLC (NuScale) has developed a small modular reactor that supports operation of up to 12 NuScale Power Modules (NPMs) at a specific site. Each NPM is an advanced, light- water, integral pressurized water reactor (PWR) that uses a high-pressure containment vessel (CNV) immersed in a reactor pool coupled with simple, redundant, passive safety-related systems. The design ensures safe plant shutdown and cooldown in the event of a loss-of- coolant accident (LOCA). Each NPM has an independent nuclear steam supply system (NSSS) that includes a nuclear core, helical-coil steam generator (SG), integral pressurizer, and a compact, high-pressure steel CNV that contains the NSSS. The secondary system includes a traditional steam-power conversion system including a steam turbine generator, condenser, and feedwater system. The integral PWR design eliminates most of the reactor piping found on conventional PWRs, thereby reducing the possibility of a pipe rupture that would result in a LOCA. Piping in the NPM containment that potentially can break is limited to the reactor coolant system (RCS) injection line, RCS discharge line, pressurizer spray supply line, and pressurizer high point vent line. The RCS injection line is supplied by the chemical and volume control system (CVCS) and the discharge line returns to the CVCS. The NPM is designed to reduce the consequences of design basis LOCAs by using redundant, simplified, passive safety-related systems that eliminate the need for emergency core cooling system (ECCS) pumps, accumulators, and water storage tanks found on conventional PWRs. During operation, flow through the reactor is driven by natural circulation resulting from the thermal driving head produced by the temperature difference between the core and the heat sink afforded by the SG. Natural circulation flow increases reliability by eliminating primary coolant pumps that can fail or lock up.

The purpose of this topical report is to present the NuScale evaluation model (EM) used to evaluate ECCS performance in the NPM for design basis LOCAs. This LOCA EM was developed following the guidelines in the evaluation model development and assessment process (EMDAP) of “Transient and Accident Analysis Methods,” Regulatory Guide (RG) 1.203 (Reference1), and adheres to the applicable requirements of “ECCS Evaluation Models,” 10 CFR 50 Appendix K (Reference 2) and “Acceptance Criteria for Emergency Core Cooling Systems for Light-Water Nuclear Power Reactors,” 10 CFR 50.46 (Reference 3). Multiple layers of conservatism are incorporated in the NuScale LOCA EM to ensure that a conservative analysis result is obtained. These conservatisms stem from application of the modeling requirements of 10 CFR 50 Appendix K and through a series of conservative modeling features.

The LOCA EM uses the proprietary NRELAP5 systems analysis computer code as the computational engine, derived from the Idaho National Laboratory (INL) RELAP5-3D© computer code. RELAP5-3D© was procured and as part of the procurement process commercial grade dedication was performed by NuScale to establish the baseline NRELAP5 code for development. Subsequently, features were added and changes made to NRELAP5 to address the unique aspects of the NPM design and licensing methodology. NRELAP5 includes all of the necessary models for characterization of the NPM hydrodynamics, heat transfer between structures and fluids, modeling of fuel, reactor kinetics models, and control systems. The models and correlations used by NRELAP5 have been reviewed and, where appropriate, modified for use within the NuScale LOCA methodology. Code changes for the NuScale application include new helical coil SG heat transfer and pressure drop models, core critical

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heat flux (CHF) models, and interfacial drag models for large-diameter pipes. The fuel CHF models were selected based on full-scale fuel bundle performance tests.

Validation and verification of the EM and NRELAP5 code were conducted in accordance with the EMDAP process. A phenomena identification and ranking table (PIRT), which identifies the important phenomena and processes occurring in the NPM during a LOCA event, was developed by gathering and ranking expert evaluations of phenomena that could occur in the NPM during a LOCA. Phenomena and process ranking was performed in relation to specified figures of merit (FOMs) as described by RG 1.203. The PIRT also established a knowledge ranking for each phenomenon identified. Using these FOMs, 21 phenomena were identified as important to capture in the NuScale LOCA EM.

Extensive NRELAP5 code validation was performed to ensure that the LOCA EM is applicable for all important phenomena and processes over the range encountered in the NPM LOCA. The validation suite includes many legacy separate effects tests (SETs) and integral effects tests (IETs), as well as many SETs and IETs developed and run specifically for the NPM application. The SETs run for the NPM application were performed at the Società Informazioni Esperienze Termoidrauliche (SIET) facility on a model helical coil SG, and at the Stern facility to obtain CHF data on a full-scale rod bundle test section. The IETs were performed at the Oregon State University NuScale Integral System Test -1 (NIST-1) facility, a scaled representation of the complete NPM primary and secondary systems, as well as the reactor pool.

The EMDAP requires an applicability demonstration of the NRELAP5 code and tests. A unique aspect of the demonstration provided for the NPM is the comparison of NRELAP5 simulations of LOCA events to NIST-1 test data and NRELAP5 simulation of the same LOCA event in the NPM. In the comparisons, the NPM results are scaled down to the NIST-1 size using the scaling ratios used to design the NIST-1 facility. The reasonable-to-excellent agreement obtained by these comparisons establishes the applicability of NRELAP5 to accurately predict LOCA phenomena at both the NIST-1 and NPM scales.

This topical report provides example applications of the LOCA EM in order to aid the reader’s understanding of the context of the application of the NuScale LOCA EM. These calculations are presented for break spectra that cover a range of break locations, break sizes, single failures, equipment unavailability and initial and boundary conditions. Nodalization and time-step sensitivity required by 10 CFR 50 Appendix K are also performed. The LOCA analyses demonstrate that the NPM retains sufficient water inventory in the primary system such that the core does not uncover, the fuel does not experience a CHF condition and the containment design pressure is not challenged. Peak cladding temperature (PCT) is shown to occur at the beginning of the LOCA event and cladding temperature decreases as the transient evolves. Because no fuel heat-up occurs for any design-basis LOCA, the following regulatory acceptance criteria from 10 CFR 50.46 are met:

(1) Peak cladding temperature remains below 2,200 degrees Fahrenheit (1,204 degrees Celsius).

(2) Maximum fuel oxidation is less than 0.17 times total cladding thickness before oxidation.

(3) Maximum hydrogen generation is less than 0.01 times that generated if all cladding were to react.

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(4) Coolable geometry is retained.

NuScale requests Nuclear Regulatory Commission (NRC) review and approval to use the LOCA EM described in this report for analyses of design basis LOCA events in the NPM. The NuScale LOCA EM includes the following components:

• LOCA PIRT • NRELAP5 code with NuScale-specific modifications • assessment of the NRELAP5 code against experimental data • demonstration of the applicability of the NRELAP5 code to LOCA analysis • input model of the NPM This LOCA EM uses a conservative bounding approach to analyzing LOCA transients that follows the guidance provided in RG 1.203 and satisfies the applicable requirements of 10 CFR 50 Appendix K. Results show that its application to the NPM demonstrates acceptable performance based upon the acceptance criteria of 10 CFR 50.46. The methodology in this report is also used to support other analyses including: 1) events as described in Topical Report TR-0516-49416-P, “Non-Loss of Coolant Accident Methodology,” 2) containment peak pressure analysis as described in Technical Report TR-0516-49084-P, “Containment Response Analysis Methodology,” 3) long term cooling as described in Technical Report, TR-0919-51299-P, “Long-Term Cooling Methodology,” and 3) inadvertent Opening of Reactor Pressure Vessel (RPV) valves, including ECCS valves as described in Appendix B of this report, “Evaluation Model for Inadvertent Opening of RPV Valves”.

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1.0 Introduction

1.1 Purpose

The purpose of this report is to present the NuScale evaluation model (EM) used to evaluate emergency core cooling system (ECCS) performance in the NuScale Power Module (NPM) for design-basis loss-of-coolant accidents (LOCAs). The LOCA EM follows the guidance provided in “Transient and Accident Analysis Methods,” Regulatory Guide (RG) 1.203 (Reference 1) and satisfies the applicable requirements of “ECCS Evaluation Models,” 10 CFR 50 Appendix K (Reference 2). NuScale requests U.S. Nuclear Regulatory Commission (NRC) approval to use the EM described in this report for analyses of design-basis LOCA events in the NPM.

1.2 Scope

This report summarizes the following:

• NPM design and operation • NuScale LOCA phenomena identification and ranking table (PIRT) • NRELAP5 input model for the NPM • NRELAP5 code features and modifications • assessment of NRELAP5 against separate effects tests (SETs) and integral effects tests (IETs) • applicability evaluation to determine the adequacy of NRELAP5 for NPM LOCA analyses

This report also provides LOCA analysis at several locations and over a spectrum of break sizes to demonstrate the application of the EM to the NPM design. Additionally, the results of sensitivity calculations performed in accordance with the applicable requirements of 10 CFR 50 Appendix K are summarized.

The scope of the NuScale LOCA EM is as follows:

• The EM is applicable to a nuclear power plant that follows the general description of the NuScale Power Plant design in Section 3.0. Applicability of the EM is based on the NuScale LOCA PIRT, which identifies and ranks those phenomena the EM must be qualified to model during a LOCA in an NPM. • The EM does not have restrictions concerning operating setpoints or loss of offsite- power conditions as long as the phenomena that occur during the progression of a LOCA have been identified by the PIRT process. • This topical report is not intended to provide final design values or results; rather, example values for the various evaluations are provided for illustrative purposes in order to aid the reader’s understanding of the context of the application of the NuScale LOCA EM.

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• The EM is qualified for thermal-hydraulic conditions that span normal operating conditions down to atmospheric pressure. Initially, the containment is at low absolute pressure conditions (subatmospheric). During a LOCA, the containment response depends primarily on the mass and energy release and, secondarily, on heat transfer processes on and within the containment shell. The mass and energy release does not depend on downstream (containment) conditions until the containment pressure is above atmospheric pressure. Hence, the lower limit for models and correlations used in the LOCA analysis is atmospheric pressure. • The EM requires that certain checks be made and conservative assumptions be taken when building the model. This includes the generation and application of a bounding power shape and the selection of a set of thermal-mechanical properties that bounds all times in cycle. • Application of the EM demonstrates that fuel does not experience CHF conditions, collapsed water level remains above the top of the active fuel, and containment remains intact and pressure and temperature remain below design limits. This assures that no fuel failure occurs and the acceptance criteria of 10 CFR 50.46 (Reference 3), excluding long-term cooling, are satisfied. • The EM described in this document addresses ECCS performance in the NPM up to the time when a recirculation flow is established. Recirculation flow is considered established when pressure and level in containment and the RPV approach a stable equilibrium condition {i.e., flow is recirculating through the reactor recirculation valves [RRVs]), core heat is removed by boiling in the core, and steam exits through the reactor vent valves (RVVs). This EM does not assess radiological impacts, boron precipitation, or boron dilution. These aspects are assessed by separate methodologies. Long term cooling is addressed in the NuScale technical report, “Long Term Cooling,” TR-0916-51299 (Reference 11). • Pipe breaks inside containment are considered to be LOCA. Pipe breaks outside containment and failures in reactor pressure vessel (RPV) appurtenances, e.g., control rod drive mechanism housings and RPV nozzles and flanges, are not evaluated as part of the LOCA definition. Inadvertent opening of valves on the RPV leading to a decrease in reactor coolant system (RCS) inventory are not included in the LOCA definition per 10 CFR 50.46. However, the LOCA EM has been extended to model such transients as described in Appendix B.

1.3 Abbreviations and Definitions

Table 1-1. Abbreviations

Term Definition ABWR advanced boiling water reactor AC alternating current AOO anticipated operational occurrence BOL beginning-of-life BWR boiling water reactor CCFL counter current flow limitation

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Term Definition CFT critical flow test CHF critical heat flux CHFR critical heat flux ratio CPV cooling pool vessel CVCS chemical and volume control system CNV containment vessel DACS data acquisition and control system DC direct current DHRS decay heat removal system DSM direct substitution method DSRS design-specific review standard ECCS emergency core cooling system EM evaluation model EMDAP evaluation model development and assessment process EOL end-of-life FLECHT full length emergency cooling heat transfer FOM figure of merit FWIV feedwater isolation valve GDF general design framework HBM heat balance method HPCF high pressure core flooder system (for BWRs and ABWRs) HPSI high pressure safety injection (for conventional PWRs) HTFS heat transfer and fluid flow service HTP heat transfer plate H2TS hierarchical two-tiered scaling IAB inadvertent actuation block ID inner diameter IET integral effects test INL Idaho National Laboratory JIT jet impingement test KATHY Karlstein Thermal-Hydraulic test facility L/D length-to-diameter LOCA loss-of-coolant accident LP lower plenum LPFL low pressure core flooder system (for ABWRs) MASLWR Multi-Application Small Light Water Reactor MCHFR minimum critical heat flux ratio MSIV main steam isolation valve MPS module protection system NIST-1 NuScale Integral System Test -1 NRC U.S. Nuclear Regulatory Commission NSSS nuclear steam supply system NPM NuScale Power Module PCT peak cladding temperature

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Term Definition PIRT phenomena identification and ranking table PWR pressurized water reactor QAPD Quality Assurance Program Description RCIC reactor core isolation cooling system (for BWRs and ABWRs) RCS reactor coolant system RG Regulatory Guide RHR residual heat removal system (conventional plants) RPV reactor pressure vessel RRV reactor recirculation valve RSV reactor safety valve RVV reactor vent valve SET separate effects test SG steam generator SIET Società Informazioni Esperienze Termoidrauliche SRV safety relief valve SSC Structures, Systems and Components TAF top of active fuel UCP upper core plate

Table 1-2. Definitions

Term Definition Figure of merit A parameter selected to characterize the plant accident response. “Excellent” agreement One of the acceptance criteria defined in RG 1.203. “Excellent” agreement applies when the code exhibits no deficiencies in modeling a given behavior. Major and minor phenomena and trends are correctly predicted. The calculated results are judged to agree closely with the data. The calculation will, with few exceptions, lay within the specified or inferred uncertainty bands of the data. The code may be used with confidence in similar applications. “Reasonable” agreement One of the acceptance criteria defined in RG 1.203. “Reasonable” agreement applies when the code exhibits minor deficiencies. Overall, the code provides an acceptable prediction. All major trends and phenomena are correctly predicted. Differences between calculation and data are greater than deemed necessary for excellent agreement. The calculation will frequently lie outside but near the specified or inferred uncertainty bands of the data. However, the correct conclusions about trends and phenomena would be reached if the code was used in similar applications.

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Term Definition “Minimal” agreement One of the acceptance criteria defined in RG 1.203. “Minimal” agreement applies when the code exhibits significant deficiencies. Overall, the code provides a prediction that is only conditionally acceptable. Some major trends or phenomena are not predicted correctly and some calculated values lie considerably outside the specified or inferred uncertainty bands of the data. Incorrect conclusions about trends and phenomena may be reached if the code were to be used in similar applications and an appropriate warning needs to be issued to users. Selected code models and facility model noding need to be reviewed, modified, and assessed before the code can be used with confidence in similar applications.

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2.0 Background

This topical report provides a description of the NuScale LOCA EM, developed following the guidelines in the EMDAP of RG 1.203.

Six basic principles are identified in RG 1.203 as important in the process of developing and assessing an EM. Four of the principles (corresponding to the 20 steps identified in the EMDAP process) are addressed in this report. They include

• determining the requirements for the EM. • developing an assessment base consistent with the determined requirements. • developing the EM. • assessing the adequacy of the EM.

The remaining principles related to establishing an appropriate quality assurance program and providing comprehensive, accurate, up-to-date documentation) are addressed outside this report as part of “NuScale Topical Report: Quality Assurance Program Description for the NuScale Power Plant,” NP-TR-1010-859-NP (Reference 4).

The NuScale LOCA EM specifically addresses the application of the EM to the NPM and how the EM meets the applicable requirements of 10 CFR 50 Appendix K. This report also demonstrates how the NuScale LOCA EM can be applied to evaluate ECCS performance to meet 10 CFR 50.46 acceptance criteria.

This EM uses the NRELAP5 code that was developed from the Idaho National Laboratory (INL) RELAP5-3D© computer code. This report discusses the code modifications and modeling requirements needed to address the unique features and phenomena of the NPM design, as well as those required to comply with the applicable requirements of 10 CFR 50 Appendix K.

The EM developed in this report is consistent with the applicable TMI Action Items (Reference 5) as described in the Design-Specific Review Standard for NuScale, Section 15.6.5 (Reference 6).

2.1 Loss-of-Coolant Accident Evaluation Model Roadmap

Figure 2-1 shows various elements of the EMDAP as defined in RG 1.203 and provides a roadmap that relates the sections of this report to the elements and steps of the EMDAP. The EMDAP establishes the adequacy of a methodology for evaluating complex events that are postulated to occur in nuclear power plant systems. The EMDAP described here has been developed for analyzing postulated LOCAs in the NPM.

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Element 1 Establish Requirements for Evaluation Model Capability

1. Specify analysis purpose, transient class and power plant class 2. Specify figures of merit 3. Identify systems, components, phases geometries, fields and processes that should be modeled 4. Identify and Rank phenomena and processes

Element 2 Element 3 Develop Assessment Base Develop Evaluation Model 5. Specify objectives for assessment base 6. Perform scaling analyses and identify 10. Establish EM development plan similarity criteria 11. Establish EM structure 7. Identify existing data and/or perform IETs 12. Develop or incorporate closure models and SETs to complete the database 8. Evaluate effects of IET distortions and SET scaleup capability 9. Determine experimental uncertainties

Element 4 Assess Evaluation Model Adequacy

Closure Relations (Bottom-up) Integrated EM (Top-down) 13. Determine model pedigree and applicability to 16. Determine capability of field equations and simulate physical processes numeric solutions to represent processes and 14. Prepare input and perform calculations to phenomena and ability of numeric solutions to assess model fidelity or accuracy approximate equation set 15. Assess scalability of models 17. Determine applicability of EM to simulate system components 18. Prepare input and perform calculations to assess system interactions and global capability 19. Assess scalability of integrated calculations and data for distortions.

20. Determine EM bases and uncertainties

No Yes Return to appropriate Adequacy Decision Perform plant elements, make and event analyses Does code meet adequacy standard? assess corrections.

Figure 2-1. Evaluation model development and assessment process

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Table 2-1. Evaluation model development and assessment process steps and the associated sections in this document

EMDAP Description EM Section Step Element 1, Establish Requirements for Evaluation Model Capability The purpose of the LOCA EM is described in Section 1.1. Section 2.0 briefly describes the background of the process followed to develop the LOCA EM and the principal software used. Specify analysis purpose, Section 3.0 provides an overview of the NPM design and 1 transient class, and power operation. This includes the safety-related systems, the system plant class. logic, and operational phases that could occur in the NPM. The regulatory requirements with which the EM is designed to comply are described in Section 2.2. Specify figures of merit Section 4.3 discusses the FOMs which are used for the 2 (FOMs). development of the NuScale LOCA PIRT. Identify systems, components, phases, Systems, components, phases, and processes are identified as a 3 geometries, fields, and part of the NuScale LOCA PIRT discussed in Section 4.0. processes that should be modeled. Identify and rank Section 4.0 summarizes the PIRT that has been established for 4 phenomena and this EM. processes. Element 2, Develop Assessment Base Specify objectives for Section 7.0 describes objectives of the benchmarks selected for 5 assessment base. the assessment of NRELAP5 against SETs and IETs. A scaling analysis has been performed for the NPM based on the Perform scaling analysis NuScale Integral System Test -1 (NIST-1) facility. The results of 6 and identify similarity the scaling analysis are discussed in Section 8.3.2 to address the criteria. EM applicability to the NPM LOCA analysis. Identify existing data Sections 7.2 through 7.5 provide the results of the NRELAP5 and/or perform IETs and validation against the SETs and IETs. In Section 8.0 these results 7 SETs to complete are evaluated relative to NRELAP5 modeling of the high-ranked database. phenomena identified in the NuScale LOCA PIRT. Evaluate effects of IET The SET scale-up capability is evaluated in Section 8.2. NIST-1 8 distortions and SET scale- IET distortions are evaluated in Section 8.3. These results justify up capability. the applicability of the EM to NPM LOCA analysis. Determine experimental Section 7 covers experimental uncertainties for NRELAP5 9 uncertainties. assessments against the SETs and IETs. Element 3, Develop Evaluation Model The NRELAP5 development plan includes programming Establish EM development 10 standards and procedures, quality assurance procedures, and plan. configuration control, which are summarized in Section 6.1.

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EMDAP Description EM Section Step The final structure of the LOCA EM is described in Section 5.0. 11 Establish EM structure. The NRELAP5 code description and new model features are discussed in Section 6.0. A full description of the closure models and the associated equations used in the LOCA EM is provided in the NRELAP5 Develop or incorporate theory and users manuals. Section 6.2 provides a summary of 12 closure models. NRELAP5 models and correlations. The applicability evaluation in Section 8.0 also provides further discussion of the NRELAP5 code models and correlations. Element 4, Assess Evaluation Model Adequacy Closure Relations (Bottom-up) Determine model pedigree Bottom-up assessments presented in Section 8.2 include and applicability to discussion of pedigree and applicability of dominant NRELAP5 13 simulate physical models and correlations that are essential to simulate high- processes. ranked PIRT phenomena. Sections 7.2 through 7.5 summarize the results of comparison of Prepare input and perform NRELAP5 against the selected SETs and IETs, including calculations to assess 14 evaluation of code fidelity and accuracy. These results are model fidelity and considered in Section 8.2 to address the applicability of the EM to accuracy. NPM LOCA analysis. Section 8.2 includes discussion of scalability of dominant Assess scalability of 15 NRELAP5 models and correlations that are essential to simulate models. high-ranked PIRT phenomena. Element 4, Assess Evaluation Model Adequacy Integrated EM (Top-down) Determine capability of field equations and NRELAP5 field equations and the numeric solution scheme are 16 numeric solutions to discussed in Section 6.2 and evaluated for their applicability to represent processes and NPM LOCA in Section 8.0. phenomena. Determine applicability of The applicability of the EM to simulate the NPM system and 17 EM to simulate system components is demonstrated by assessment of NRELAP5 components. against NuScale design-specific SETs and IETs in Section 8.3.1. Prepare input and perform Section 7.0 summarizes the results of the assessment of calculations to assess NRELAP5 against NIST-1 IET data. These results are considered 18 system interactions and in Section 8.3 to address the applicability of the EM to NPM global capability. LOCA analysis. Section 8.3 provides an evaluation of scaling distortions between Assess scalability of the NIST-1 IET data and the NPM design. The scalability of EM 19 integrated calculations and to represent NPM LOCA phenomena and processes is data for distortions. presented. Determine EM biases and This step is not required per RG 1.203 for safety analyses that 20 uncertainties. implement 10 CFR 50 Appendix K.

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2.2 Regulatory Requirements

This section discusses the regulatory acceptance criteria for ECCS performance and the manner in which they are satisfied by application of the NuScale LOCA EM.

2.2.1 10 CFR 50.46 Loss-of-Coolant Accident Acceptance Criteria

10 CFR 50.46 requires that light water nuclear reactors fueled with uranium oxide pellets within cylindrical zircaloy cladding be provided with an ECCS that is designed in such a way that their calculated core cooling performance after a postulated LOCA conforms to certain criteria specified in 10 CFR 50.46(b). The five acceptance criteria are the following:

1. The calculated maximum fuel element cladding temperature shall not exceed 2200 degrees Fahrenheit (1,204 degrees Celsius).

2. The calculated total oxidation of the cladding shall nowhere exceed 0.17 times the total cladding thickness before oxidation.

3. The calculated total amount of hydrogen generated from the chemical reaction of the cladding with water or steam shall not exceed 0.01 times the hypothetical amount that would be generated if all the metal in the cladding cylinders surrounding the fuel, excluding the cladding surrounding the plenum volume, were to react.

4. Calculated changes in core geometry shall be such that the core remains amenable to cooling.

5. After any calculated successful initial operation of the ECCS, the calculated core temperature shall be maintained at an acceptably low value and decay heat shall be removed for the extended period of time required by the long-lived radioactivity remaining in the core.

The NuScale LOCA EM addresses the first four criteria as described in Section 2.2.2. The EM described in this document addresses ECCS performance in the NPM up to the time when a recirculation flow is established, pressures and levels in containment and the RPV approach a stable equilibrium condition (i.e., flow is recirculating in through the RRVs), core heat is removed by boiling in the core, and steam exits through the RVVs.

2.2.2 NuScale Loss-of-Coolant Accident Evaluation Model Acceptance Criteria

The NPM is designed so that there is no core uncovery or heatup for a design-basis LOCA. As a result, peak cladding temperature (PCT) will be well within the acceptance criterion of 2,200 degrees Fahrenheit (1,204 degrees Celsius). The parameters of interest are the collapsed liquid water level above the top of active fuel (TAF) and minimum critical heat flux ratio (MCHFR). These two criteria are more sensitive than PCT in the NPM design. Maintaining primary inventory and ensuring the core does not go into post-critical heat flux (CHF) heat transfer ensures that the 10 CFR 50.46(b) limitations for PCT, oxidation, and hydrogen production are protected.

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There is no oxidation of the cladding as a result of a LOCA. There is no hydrogen generated from the chemical reaction of the cladding with water or steam because fuel temperatures are not high enough to initiate this chemical reaction. There are no changes in core geometry resulting from a LOCA that would prevent the core from being amenable to cooling. Therefore, the first four acceptance criteria are met when the collapsed liquid level is above the top of the active fuel and MCHFR is greater than the analysis limit for the entire time period covered by this EM (see Section 7.3.6).

The fifth criterion is also met during the shorter period this EM addresses. The longer- term evaluation for the fifth criteria is addressed by other NuScale methodologies (Reference 11).

In summary, the NuScale LOCA EM acceptance criteria are:

1. Collapsed liquid level (see Section 5.1.2.6) remains above the top of the active fuel, and

2. MCHFR is greater than analysis limit of 1.29 (see Section 7.3.6)

2.2.3 10 CFR 50 Appendix K

The ECCS performance is calculated in conformance with the required and acceptable features of ECCS EMs specified in 10 CFR 50 Appendix K, and is calculated for a number of cases to provide assurance that the most severe postulated LOCAs are identified. 10 CFR 50.46 provides two options for an acceptable LOCA EM. Paragraph 50.46(a)(1)(i) allows for a best-estimate approach to be followed and Paragraph 50.46.(a)(ii) allows for the conservative deterministic approach detailed in 10 CFR 50 Appendix K. In view of the large safety margins in the NPM, the deterministic bounding approach in Paragraph 50.46(a)(1)(ii) is used by NuScale.

The NPM is designed to reduce the consequences of design-basis LOCAs compared to existing light water reactors for which 10 CFR 50 Appendix K was developed. Consequently, many of the phenomena that are the subject of 10 CFR 50 Appendix K requirements are not encountered in design-basis NPM LOCAs in the NPM. That is, certain phenomena have been designed out of the NPM and, therefore, a number of requirements are satisfied by design rather than by analysis. Examples of phenomena and processes that can occur during a typical pressurized water reactor (PWR) LOCA that do not occur during an NPM LOCA include:

• loop seal clearing • pump coastdown • two-phase pump performance • entry of significant amounts of non-condensable gases into the system • core uncovery • core refilling

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• core reflooding • cladding swelling and rupture • metal-water reaction • post-CHF heat transfer • cladding rewet • ECCS bypass

Hence, only a subset of the phenomena that are addressed in 10 CFR 50 Appendix K is encountered in the design-basis NPM LOCAs and thus relevant to the NuScale LOCA EM. Table 2-2 lists each required and acceptable feature of the EM specified in 10 CFR 50 Appendix K and describes the manner in which the NuScale LOCA EM addresses each feature. The NuScale LOCA EM includes model features required by Appendix K that are relevant to the NPM. Features required by 10 CFR 50 Appendix K that are not relevant to the NuScale LOCA EM are identified in Table 2-2 as either “satisfied by design” or “excluded from model.”

A feature “satisfied by design” means that a 10 CFR 50 Appendix K required feature is expressly or impliedly conditional on the presence of process or phenomena in the design or analysis. Because such process or phenomena does not exist for the NuScale design, the required feature is not applicable and not included in the NuScale LOCA EM. For example, there are no reactor coolant pumps in the NPM. Therefore the phenomena that are the subject of 10 CFR 50 Appendix K Requirement I.C.6 “Pump Modeling” are not encountered because of the design of the NPM, and thus the required model features are “satisfied by design.”

A feature “excluded” from the EM means that 10 CFR 50 Appendix K directly requires the feature, without condition on the presence of a process or phenomena, but that the feature is not relevant to the NuScale LOCA EM. Table 2-2 technically justifies the exclusion of such feature from the model. However, an applicant or licensee referencing this report will be required to address regulatory compliance with 10 CFR 50.46 and 10 CFR 50 Appendix K (e.g., by seeking an exemption from that required feature). Similarly, an “acceptable alternative” model feature is technically justified by Table 2-2, but does not strictly meet the 10 CFR 50 Appendix K required feature, and thus an applicant or licensee referencing this report will be required to address regulatory compliance. Historically, RELAP5 has been applied to evaluate post-CHF fuel conditions for events in LWRs. While these features have been retained in NRELAP5, the application of the LOCA EM to predict fuel temperature response is limited to pre-CHF heat transfer regimes.

In the NuScale LOCA EM, applicable closure models or correlations required by 10 CFR 50 Appendix K are used. The NuScale LOCA EM also uses appropriate closure models or correlations in addition to those required in 10 CFR 50 Appendix K. All closure models and correlations are verified and validated for use within their range of applicability.

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Table 2-2. 10 CFR 50 Appendix K required and acceptable features compliance

10 CFR 50 Appendix K Required and NuScale LOCA EM Acceptable Feature I.A Sources of heat during the LOCA:

For the heat sources listed in paragraphs I.A.1 to 4 of this appendix, it must be assumed that the reactor has been operating continuously at a The initial power level is set at 102 percent of power level at least 1.02 times the licensed power rated power. The maximum radial peaking factor is level (to allow for instrumentation error) with the used in the hot assembly to bound all possible maximum peaking factor allowed by the technical power peaking. (See Section 9.6.6). Sensitivity specifications. An assumed power level lower than calculations were performed with different axial the level specified in this paragraph (but not less power shapes that bound maximum axial power than the licensed power level) may be used peaking. provided the proposed alternative value has been demonstrated to account for uncertainties due to Further discussion on core power distribution is power level instrumentation error. A range of provided in Section 5.1.2.2.3. power distribution shapes and peaking factors representing power distributions that may occur over the core lifetime must be studied. The Therefore, the required features of I.A are selected combination of power distribution shape included in the NuScale LOCA EM. and peaking factor should be the one that results in the most severe calculated consequences for the spectrum of postulated breaks and single failures that are analyzed.

I.A.1 The Initial Stored Energy in the Fuel: Based on the burn-up dependent fuel performance analysis, it was determined that choosing end-of-life (EOL) fuel thermal The steady-state temperature distribution and conductivity and beginning-of-life (BOL) stored energy in the fuel before the hypothetical volumetric heat capacity and fuel-cladding gap accident shall be calculated for the burn-up that conductance maximizes the initial stored energy yields the highest calculated cladding temperature in the fuel. An additional 15 percent bias is (or, optionally, the highest calculated stored applied to both volumetric heat capacity and energy.) To accomplish this, the thermal thermal conductivity to maximize the initial stored conductivity of the UO2 shall be evaluated as a energy. function of burn-up and temperature, taking into consideration differences in initial density, and the Further discussion on selection of fuel rod thermal conductance of the gap between the UO2 and the cladding shall be evaluated as a function mechanical property input is provided in Section of the burn-up, taking into consideration fuel 5.1.2.2.4. densification and expansion, the composition and pressure of the gases within the fuel rod, the initial Therefore, the required features of I.A.1 are cold gap dimension with its tolerances, and included in the NuScale LOCA EM. cladding creep.

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10 CFR 50 Appendix K Required and NuScale LOCA EM Acceptable Feature A point kinetics model is used to calculate fission I.A.2 Fission Heat: power. Credit is taken for reactor trip. A conservative control rod insertion curve is used along with a minimum rod worth and conservative Fission heat shall be calculated using reactivity delay in initiation of rod insertion. The most and reactor kinetics. Shutdown reactivities reactive control rod is assumed to be stuck out of resulting from temperatures and voids shall be the core. Doppler and moderator density given their minimum plausible values, including coefficients are calculated conservatively (Section allowance for uncertainties, for the range of power 5.1.2.2.5). distribution shapes and peaking factors indicated to be studied above. Rod trip and insertion may be assumed if they are calculated to occur. Therefore, the required features of I.A.2 are included in the NuScale LOCA EM.

I.A.3 Decay of Actinides: The 1979 ANS actinide decay heat standard is applied which includes the decay of neptunium

and plutonium (Sections 5.1.2.2.5). The heat from the radioactive decay of actinides, including neptunium and plutonium generated during operation, as well as isotopes of uranium, The actinide decay heat assumes infinite shall be calculated in accordance with fuel cycle operating time to maximize actinide concentration. calculations and known radioactive properties. The This assumption results in the highest calculated actinide decay heat chosen shall be that fuel temperature during the LOCA. appropriate for the time in the fuel cycle that yields the highest calculated fuel temperature during the Therefore, the required features of I.A.3 are LOCA. included in the NuScale LOCA EM. I.A.4 Fission Product Decay:

The heat generation rates from radioactive decay of fission products shall be assumed to be equal to 1.2 times the values for infinite operating time in The 1973 ANS decay heat standard (Reference the ANS Standard (Proposed American Nuclear 44) is used with a 20 percent uncertainty added to Society Standards--"Decay Energy Release Rates the base value. A bounding form of the 1973 ANS Following Shutdown of Uranium-Fueled Thermal standard in NRELAP5 meets the intent of the 10 Reactors." Approved by Subcommittee ANS-5, CFR 50 Appendix K requirement (Section ANS Standards Committee, October 1971). This 5.1.2.2.5). standard has been approved for incorporation by reference by the Director of the Federal Register. A copy of the standard is available for inspection Therefore, the NuScale LOCA EM includes an at the NRC Library, 11545 Rockville Pike, acceptable alternative to the requirement of I.A.4. Rockville, Maryland 20852-2738. The fraction of the locally generated gamma energy that is deposited in the fuel (including the cladding) may be different from 1.0; the value used shall be justified by a suitable calculation.

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10 CFR 50 Appendix K Required and NuScale LOCA EM Acceptable Feature I.A.5 Metal-Water Reaction Rate:

The rate of energy release, hydrogen generation, and cladding oxidation from the metal-water reaction shall be calculated using the Baker-Just equation (Baker, L., Just, L.C., "Studies of Metal Calculated cladding temperatures for design basis Water Reactions at High Temperatures, III. LOCAs are well below the level where cladding Experimental and Theoretical Studies of the oxidation occurs on a time scale of a LOCA event Zirconium-Water Reaction," ANL-6548, page 7, for the NPM (see the results of LOCA break May 1962). This publication has been approved for spectrum calculations in Section 9.0). Therefore, incorporation by reference by the Director of the this requirement is not relevant to the NuScale Federal Register. A copy of the publication is design, which precludes fuel temperature reaching available for inspection at the NRC Library, 11545 CHF and any significant fuel cladding heatup. For Rockville Pike, Two White Flint North, Rockville, the NuScale LOCA EM, core coverage and an Maryland 20852-2738. The reaction shall be MCHFR greater than the analysis limit (see assumed not to be steam limited. For rods whose Section 7.3.6) precludes the occurrence of cladding is calculated to rupture during the LOCA, cladding oxidation. the inside of the cladding shall be assumed to react after the rupture. The calculation of the Therefore, the required features of I.A.5 are reaction rate on the inside of the cladding shall excluded from the NuScale LOCA EM. also follow the Baker-Just equation, starting at the time when the cladding is calculated to rupture, and extending around the cladding inner circumference and axially no less than 1.5 inches each way from the location of the rupture, with the reaction assumed not to be steam limited. The NRELAP5 plant model explicitly represents all major reactor internal heat structures. Heat structures are also included for the primary and I.A.6 Reactor Internals Heat Transfer: secondary system pressure boundary materials. See Section 5.1.2 for details of the internal heat Heat transfer from piping, vessel walls, and non- structures represented in the NuScale LOCA EM. fuel internal hardware shall be taken into account. Therefore, the required features of I.A.6 are included in the NuScale LOCA EM. Heat transfer through the steam generator (SG) I.A.7 Pressurized Water Reactor Primary-to- tubes is included in the EM. The model is Secondary Heat Transfer: validated using experimental data from Società Italiana Esperienze Termoidrauliche (SIET) tests (see Section 7.4) and NIST-1 tests (see Section Heat transferred between primary and secondary 7.5). systems through heat exchangers (steam generators) shall be taken into account. (Not applicable to boiling water reactors (BWRs).) Therefore, the required features of I.A.7 are included in the NuScale LOCA EM.

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10 CFR 50 Appendix K Required and NuScale LOCA EM Acceptable Feature I.B Swelling and Rupture of the Cladding and Fuel Rod Thermal Parameters:

Each evaluation model shall include a provision for predicting cladding swelling and rupture from Calculated cladding temperatures for design basis consideration of the axial temperature distribution LOCAs in the NPM are well below the threshold of the cladding and from the difference in pressure for cladding swelling and rupture (see the results between the inside and outside of the cladding, of LOCA break spectrum calculations in Section both as functions of time. To be acceptable the 9.0). Peak cladding temperatures in the NPM swelling and rupture calculations shall be based occur at steady state normal operation. Because on applicable data in such a way that the degree swelling and rupture do not occur during normal of swelling and incidence of rupture are not operation, they will not occur in a NPM LOCA underestimated. The degree of swelling and event. Therefore, this requirement is not relevant rupture shall be taken into account in calculations for the NuScale LOCA EM as core coverage of gap conductance, cladding oxidation and precludes the occurrence of cladding swelling and embrittlement, and hydrogen generation. rupture.

The calculations of fuel and cladding temperatures Therefore, the required features of I.B are as a function of time shall use values for gap excluded from the NuScale LOCA EM. conductance and other thermal parameters as functions of temperature and other applicable time-dependent variables. The gap conductance shall be varied in accordance with changes in gap dimensions and any other applicable variables. I.C Blowdown Phenomena A complete spectrum of break sizes and locations is analyzed up to the largest penetrations in the RPV including the double-ended guillotine break I.C.1.a Break Characteristics and Flow: where appropriate. The size of the pipes precludes the impact of longitudinal split breaks in In analyses of hypothetical LOCAs, a spectrum of the NPM design. Therefore, the requirement for possible pipe breaks shall be considered. This analyzing the effect of longitudinal split break is spectrum shall include instantaneous double- not relevant to the NuScale LOCA EM. ended breaks ranging in cross-sectional area up to and including that of the largest pipe in the primary Further discussion of break spectrum analysis is coolant system. The analysis shall also include the provided in Section 5.4. The break spectrum effects of longitudinal splits in the largest pipes, calculation results are available in Section 9.0. with the split area equal to the cross-sectional area of the pipe. Therefore, the required features of I.C.1.a are included or satisfied by design in the NuScale LOCA EM.

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10 CFR 50 Appendix K Required and NuScale LOCA EM Acceptable Feature The required Moody critical flow is used when the break flow is calculated to be two-phase flow and the {{ }}2(a),(c) model is used to I.C.1.b Discharge Model: calculate single-phase choked flow {{

For all times after the discharging fluid has been }}2(a),(c) For the NPM, single calculated to be two-phase in composition, the phase flow through the break may recur after the discharge rate shall be calculated by use of the transition to two-phase flow. The {{ Moody model (F.J. Moody, "Maximum Flow Rate of a Single Component, Two-Phase Mixture," Journal of Heat Transfer, Trans American Society }}2(a),(c) model is conservative for single- of Mechanical Engineers, 87, No. 1, February, phase break flow. See Section 6.6.1 for details. 1965). The calculation shall be conducted with at least three values of a discharge coefficient The range of postulated break sizes in the break applied to the postulated break area, these values analysis covers the 10 CFR 50 Appendix K spanning the range from 0.6 to 1.0. If the results required range of discharge coefficient, as indicate that the maximum cladding temperature discussed in Section 5.4. for the hypothetical accident is to be found at an even lower value of the discharge coefficient, the range of discharge coefficients shall be extended Therefore, the required features of I.C.1.b, until the maximum cladding temperatures including an acceptable alternative feature, are calculated by this variation has been achieved. included in the NuScale LOCA EM.

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10 CFR 50 Appendix K Required and NuScale LOCA EM Acceptable Feature I.C.1.c End of Blowdown. (Applies Only to Pressurized Water Reactors):

For postulated cold leg breaks, all emergency cooling water injected into the inlet lines or the reactor vessel during the bypass period shall in the calculations be subtracted from the reactor vessel calculated inventory. This may be executed in the calculation during the bypass period, or as an alternative the amount of emergency core cooling For the NuScale design, there are no cold legs water calculated to be injected during the bypass and hence no cold leg breaks. All of the coolant period may be subtracted later in the calculation that exits the break remains in the containment from the water remaining in the inlet lines, and is available to return when the RRVs are downcomer, and reactor vessel lower plenum after opened. Emergency core cooling system bypass the bypass period. This bypassing shall end in the cannot occur in the NPM, so this requirement is calculation at a time designated as the "end of not relevant to the NuScale LOCA EM. bypass," after which the expulsion or entrainment mechanisms responsible for the bypassing are calculated not to be effective. The end-of-bypass Therefore, the required features of I.C.1.c are definition used in the calculation shall be justified satisfied by design. by a suitable combination of analysis and experimental data. Acceptable methods for defining "end of bypass" include, but are not limited to, the following: (1) Prediction of the blowdown calculation of downward flow in the downcomer for the remainder of the blowdown period; (2) Prediction of a threshold for droplet entrainment in the upward velocity, using local fluid conditions and a conservative critical Weber number. Noding sensitivity studies have been conducted to I.C.1.d Noding Near the Break and the ECCS demonstrate that the calculated conditions in the Injection Points: vicinity of the break locations, RVVs, and RRVs are reliable.

The noding in the vicinity of and including the broken or split sections of pipe and the points of The results of the noding sensitivity studies are ECCS injection shall be chosen to permit a reliable discussed in Section 9.6.1. analysis of the thermodynamic history in these regions during blowdown. Therefore, the required features of I.C.1.d are included in the NuScale LOCA EM.

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10 CFR 50 Appendix K Required and NuScale LOCA EM Acceptable Feature I.C.2 Frictional Pressure Drops:

The frictional losses in pipes and other components including the reactor core shall be calculated using models that include realistic variation of friction factor with Reynolds number, and realistic two-phase friction multipliers that Friction losses in pipes and components are have been adequately verified by comparison with calculated using Reynolds number-dependent experimental data, or models that prove at least friction factors. The NRELAP5 wall friction model equally conservative with respect to maximum is based on a two-phase multiplier approach (see cladding temperature calculated during the Section 6.2.4). The models used in NRELAP5 hypothetical accident. The modified Baroczy have been validated for the range of conditions correlation (Baroczy, C. J., "A Systematic encountered in design-basis LOCAs as shown by Correlation for Two-Phase Pressure Drop," Chem. assessment against SETs and IETs in Section Enging. Prog. Symp. Series, No. 64, Vol. 62, 7.0. 1965) or a combination of the Thom correlation (Thom, J.R.S., "Prediction of Pressure Drop During Forced Circulation Boiling of Water," Int. J. Therefore, the required features of I.C.2 are of Heat & Mass Transfer, 7, 709-724, 1964) for included in the NuScale LOCA EM. pressures equal to or greater than 250 psia and the Martinelli-Nelson correlation (Martinelli, R. C. Nelson, D.B., "Prediction of Pressure Drop During Forced Circulation Boiling of Water," Transactions of ASME, 695-702, 1948) for pressures lower than 250 psia is acceptable as a basis for calculating realistic two-phase friction multipliers. I.C.3 Momentum Equation: All of the momentum equation effects required by The following effects shall be taken into account in Section I.C.3 are included in NRELAP5 (see the conservation of momentum equation: (1) Sections 6.2.1 and 6.2.4). Benchmarks for the temporal change of momentum, (2) momentum NIST-1 facility and other assessments show that convection, (3) area change momentum flux, (4) simulations made by NRELAP5 are acceptable, momentum change due to compressibility, (5) based on reasonable-to- excellent agreement with pressure loss resulting from wall friction, (6) experimental data (see Section 7.0). pressure loss resulting from area change, and (7) gravitational acceleration. Any omission of one or Therefore, the required features of I.C.3 are more of these terms under stated circumstances included in the NuScale LOCA EM. shall be justified by comparative analyses or by experimental data.

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10 CFR 50 Appendix K Required and NuScale LOCA EM Acceptable Feature Two CHF correlations are used to monitor for CHF occurrence, {{ I.C.4.a (Critical Heat Flux):

}}2(a),(c) See Sections 6.10.3 and Correlations developed from appropriate steady- 6.10.4 for description of the correlations. Section state and transient-state experimental data are 7.3 describes the assessment against the acceptable for use in predicting the CHF during NuScale CHF data that bounds the range of LOCA transients. The computer programs in which LOCA parameters.. The NuScale LOCA EM these correlations are used shall contain suitable checks to ensure that the physical parameters are checks to ensure that the physical parameters are within the range of parameters specified for use of within the range of parameters specified for use of the correlations. the correlations by their respective authors. Therefore, the required features of I.C.4.a are included in the NuScale LOCA EM. I.C.4.b (Critical Heat Flux): I.C.4.b identifies acceptable, but not required, EM Steady-state CHF correlations acceptable for use features. The NuScale LOCA EM includes an in LOCA transients include, but are not limited to, acceptable steady-state CHF correlation as the following: [six acceptable CHF correlations are addressed by I.C.4.a. identified in 10 CFR 50 Appendix K, I.C.4.b]. I.C.4.c (Critical Heat Flux):

Correlations of appropriate transient CHF data may be accepted for use in LOCA transient analyses if comparisons between the data and the correlations are provided to demonstrate that the I.C.4.c identifies acceptable, but not required, EM correlations predict values of CHF which allow for features. The NuScale LOCA EM does not use a uncertainty in the experimental data throughout transient CHF correlation. See I.C.4.a. the range of parameters for which the correlations are to be used. Where appropriate, the comparisons shall use statistical uncertainty analysis of the data to demonstrate the conservatism of the transient correlation. I.C.4.d (Critical Heat Flux):

I.C.4.d identifies acceptable, but not required, EM Transient CHF correlations acceptable for use in features. The NuScale LOCA EM does not use a LOCA transients include, but are not limited to, the transient CHF correlation. See I.C.4.a. following: (GE transient CHF correlation is listed in 10 CFR 50 Appendix K, I.C.4.d.)

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10 CFR 50 Appendix K Required and NuScale LOCA EM Acceptable Feature I.C.4.e (Critical Heat Flux):

The break analysis in Section 9.0 demonstrates After CHF is first predicted at an axial fuel rod that CHF does not occur in the NPM for LOCAs. location during blowdown, the calculation shall not Heat transfer beyond CHF is not a phenomenon use nucleate boiling heat transfer correlations at encountered during a design-basis LOCA. that location subsequently during the blowdown even if the calculated local fluid and surface The NuScale LOCA methodology does not conditions would apparently justify the calculate heat transfer beyond CHF in the core. reestablishment of nucleate boiling. Heat transfer assumptions characteristic of return to nucleate Therefore, this requirement is satisfied by a design boiling (rewetting) shall be permitted when justified that has a margin to CHF for LOCA events. by the calculated local fluid and surface conditions during the reflood portion of a LOCA. I.C.5.a (Post-CHF Heat Transfer Correlations):

Correlations of heat transfer from the fuel cladding The break analysis in Section 9.0 demonstrates to the surrounding fluid in the post-CHF regimes of that CHF does not occur in the NPM for LOCAs. transition and film boiling shall be compared to Heat transfer beyond CHF is not a phenomenon applicable steady-state and transient-state data encountered during a design-basis LOCA. using statistical correlation and uncertainty Therefore, this requirement is not relevant to the analyses. Such comparison shall demonstrate that NuScale LOCA EM. the correlations predict values of heat transfer coefficient equal to or less than the mean value of the applicable experimental heat transfer data Therefore, the required features of I.C.5.a are throughout the range of parameters for which the excluded from the NuScale LOCA EM. correlations are to be used. The comparisons shall quantify the relation of the correlations to the statistical uncertainty of the applicable data.

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10 CFR 50 Appendix K Required and NuScale LOCA EM Acceptable Feature I.C.5.b (Post-CHF Heat Transfer Correlations):

The Groeneveld flow film boiling correlation (equation 5.7 of D.C. Groeneveld, "An Investigation of Heat Transfer in the Liquid Deficient Regime," AECL-3281, revised December 1969) and the Westinghouse correlation of steady- state transition boiling ("Proprietary Redirect/Rebuttal Testimony of Westinghouse Electric Corporation," USNRC Docket RM-50-1, page 25-1, October 26, 1972) are acceptable for use in the post-CHF boiling regimes. In addition, the transition boiling correlation of McDonough, Milich, and King (J.B. McDonough, W. Milich, E.C. King, "An Experimental Study of Partial Film Boiling Region with Water at Elevated Pressures in a Round Vertical Tube," Chemical Engineering I.C.5.b identifies acceptable, but not required, EM Progress Symposium Series, Vol. 57, No. 32, features. The NuScale LOCA methodology does pages 197-208, (1961) is suitable for use between not calculate heat transfer beyond CHF in the nucleate and film boiling. Use of all these core. Therefore, these acceptable correlations are correlations is restricted as follows: not relevant to the EM. See I.C.5.a.

(1) The Groeneveld correlation shall not be used in the region near its low-pressure singularity,

(2) The first term (nucleate) of the Westinghouse correlation and the entire McDonough, Milich, and King correlation shall not be used during the blowdown after the temperature difference between the cladding and the saturated fluid first exceeds 300°F,

(3) Transition boiling heat transfer shall not be reapplied for the remainder of the LOCA blowdown, even if the cladding superheat returns below 300°F, except for the reflood portion of the LOCA when justified by the calculated local fluid and surface conditions.

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10 CFR 50 Appendix K Required and NuScale LOCA EM Acceptable Feature I.C.5.c (Post-CHF Heat Transfer Correlations):

Evaluation models approved after October 17, 1988, which make use of the Dougall-Rohsenow flow film boiling correlation (R.S. Dougall and W.M. Rohsenow, "Film Boiling on the Inside of Vertical Tubes with Upward Flow of Fluid at Low Qualities," MIT Report Number 9079 26, Cambridge, Massachusetts, September 1963) may not use this correlation under conditions where nonconservative predictions of heat transfer result. Evaluation models that make use of the I.C.5.c identifies acceptable, but not required, EM Dougall-Rohsenow correlation and were approved features. The NuScale LOCA methodology does prior to October 17, 1988, continue to be not calculate heat transfer beyond CHF in the acceptable until a change is made to, or an error is core. Therefore, these acceptable correlations are corrected in, the evaluation model that results in a not relevant to the EM. See I.C.5.a. significant reduction in the overall conservatism in the evaluation model. At that time continued use of the Dougall-Rohsenow correlation under conditions where nonconservative predictions of heat transfer result will no longer be acceptable. For this purpose, a significant reduction in the overall conservatism in the evaluation model would be a reduction in the calculated peak fuel cladding temperature of at least 50°F from that which would have been calculated on October 17, 1988, due either to individual changes or error corrections or the net effect of an accumulation of changes or error corrections. I.C.6 Pump Modeling:

The characteristics of rotating primary system pumps (axial flow, turbine, or centrifugal) shall be derived from a dynamic model that includes There are no primary system coolant pumps, so momentum transfer between the fluid and the the requirements related to pump models are not rotating member, with variable pump speed as a relevant to the NuScale LOCA EM, as shown in function of time. The pump model resistance used Section 3.0. for analysis should be justified. The pump model for the two-phase region shall be verified by applicable two-phase pump performance data. For Therefore, the required features of I.C.6 are BWRs after saturation is calculated at the pump satisfied by design. suction, the pump head may be assumed to vary linearly with quality, going to zero for one percent quality at the pump suction, so long as the analysis shows that core flow stops before the quality at pump suction reaches one percent.

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10 CFR 50 Appendix K Required and NuScale LOCA EM Acceptable Feature The core is represented by three non-interacting channels: hot channel represents hot assembly, average channel represents rest of the core assemblies, and total core bypass. The assumption of no crossflow between the core I.C.7.a Core Flow Distribution During Blowdown. regions results in conservative flow distribution. (Applies only to pressurized water reactors): Cladding swelling or rupture does not occur The flow rate through the hot region of the core because the fuel does not encounter a CHF event during blowdown shall be calculated as a function and because the core remains covered throughout of time. For the purpose of these calculations the the LOCA event. Therefore, cross flows will not be hot region chosen shall not be greater than the impacted by geometrical changes in the fuel size of one fuel assembly. Calculations of average during the transient. flow and flow in the hot region shall take into account cross flow between regions and any flow Due to the mild nature of natural circulation flow blockage calculated to occur during blowdown as during blowdown, rapid oscillations during the a result of cladding swelling or rupture. The LOCA transient with a period less than 0.1 second calculated flow shall be smoothed to eliminate any do not occur. Therefore, smoothing is not calculated rapid oscillations (period less than 0.1 necessary. seconds).

Therefore, the required features of I.C.7.a are included in the NuScale LOCA EM, except that cross-flow is conservatively excluded from the model. I.C.7.b Core Flow Distribution During Blowdown. The intention of the I.C.7.b requirement was to (Applies only to pressurized water reactors)]: ensure that LOCA EMs that assessed the hot channel separately would use the correct thermal- hydraulic boundary conditions. A method shall be specified for determining the enthalpy to be used as input data to the hot channel heatup analysis from quantities calculated For the NuScale LOCA EM, the active core is in the blowdown analysis, consistent with the flow represented by {{ distribution calculations. }}2(a),(c) The hot channel is not analyzed in a separate code, but is included in the NPM model. (See Section 5.1.2.2 for core nodalization.)

Therefore, the required features of I.C.7.b are included in the NuScale LOCA EM.

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10 CFR 50 Appendix K Required and NuScale LOCA EM Acceptable Feature I.D Post-Blowdown Phenomena; Heat Removal by the ECCS Safety-related system single failures considered I.D.1 Single Failure Criterion: for break spectrum calculations are discussed in Section 5.4.3. An evaluation of ECCS failure modes has been performed. Sensitivity studies An analysis of possible failure modes of ECCS were conducted to determine the limiting single equipment and of their effects on ECCS failure for each type of LOCA. The results of break performance must be made. In carrying out the spectrum calculations are discussed in Section accident evaluation the combination of ECCS 9.0. subsystems assumed to be operative shall be those available after the most damaging single failure of ECCS equipment has taken place. Therefore, the required features of I.D.1 are included in the NuScale LOCA EM. The NPM containment design is intended to equilibrate RCS and containment vessel (CNV) I.D.2 Containment Pressure: pressure when ECCS has been actuated. Condensed effluent will then be returned to the RCS in natural circulation flow. Although there are The containment pressure used for evaluating no active pressure-reducing systems, the CNV is cooling effectiveness during reflood and spray immersed in the reactor pool, resulting in cooling shall not exceed a pressure calculated significant condensation and cooling of effluent conservatively for this purpose. The calculation prior to returning to the RPV. shall include the effects of operation of all installed pressure-reducing systems and processes. Therefore, the required features of I.D.2 are satisfied by design.

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10 CFR 50 Appendix K Required and NuScale LOCA EM Acceptable Feature I.D.3 Calculation of Reflood Rate for Pressurized Water Reactors:

The refilling of the reactor vessel and the time and rate of reflooding of the core shall be calculated by an acceptable model that takes into consideration the thermal and hydraulic characteristics of the core and of the reactor system. The primary system coolant pumps shall be assumed to have locked impellers if this assumption leads to the maximum calculated cladding temperature; Refilling or reflooding is not required for the otherwise the pump rotor shall be assumed to be NuScale design as in a conventional PWR, running free. The ratio of the total fluid flow at the because there is no core uncovery (see the results core exit plane to the total liquid flow at the core of LOCA break spectrum calculations in Section inlet plane (carryover fraction) shall be used to 9.0). This requirement is not relevant to the determine the core exit flow and shall be NuScale LOCA EM. determined in accordance with applicable experimental data (for example, "PWR FLECHT There are no primary system coolant pumps, so (Full Length Emergency Cooling Heat Transfer) the requirements related to pump models are Final Report," Westinghouse Report WCAP-7665, satisfied by the NuScale design. Also, there are no April 1971; "PWR Full Length Emergency Cooling accumulators, so requirements related to Heat Transfer (FLECHT) Group I Test Report," accumulator discharge are satisfied by being Westinghouse Report WCAP-7435, January 1970; designed out of the NPM. "PWR FLECHT (Full Length Emergency Cooling Heat Transfer) Group II Test Report," Westinghouse Report WCAP-7544, September 1970; "PWR FLECHT Final Report Supplement," Westinghouse Report WCAP-7931, October 1972).

The effects on reflooding rate of the compressed gas in the accumulator which is discharged following accumulator water discharge shall also be taken into account.

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10 CFR 50 Appendix K Required and NuScale LOCA EM Acceptable Feature I.D.4 Steam Interaction with Emergency Core Refilling or reflooding is not required for the Cooling Water in Pressurized Water Reactors: NuScale design as in a conventional PWR, because there is no core uncovery (see the results of LOCA break spectrum calculations in Section The thermal-hydraulic interaction between steam 9.0). Traditional concerns regarding steam and all emergency core cooling water shall be interaction with injected ECCS water are not a taken into account in calculating the core factor in the NuScale design, although the reflooding rate. During refill and reflood, the phenomenon of non-equilibrium conditions calculated steam flow in unbroken reactor coolant existing between steam and subcooled liquid does pipes shall be taken to be zero during the time that occur. For the NuScale design, such interactions accumulators are discharging water into those could occur in either the CNV or in the downcomer pipes unless experimental evidence is available when subcooled containment liquid enters from regarding the realistic thermal-hydraulic interaction the RRVs. While I.D.4 is not relevant to the between the steam and the liquid. In this case, the NuScale LOCA EM, the intent of this requirement experimental data may be used to support an is addressed by the capability of NRELAP5 to alternate assumption. model thermal non-equilibrium states and by the NPM design which minimizes these phenomena.

Therefore, the required features of I.D.4 are satisfied by design.

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10 CFR 50 Appendix K Required and NuScale LOCA EM Acceptable Feature I.D.5.a (Refill and Reflood Heat Transfer for Pressurized Water Reactors):

For reflood rates of one inch per second or higher, reflood heat transfer coefficients shall be based on applicable experimental data for unblocked cores including FLECHT results ("PWR FLECHT (Full Length Emergency Cooling Heat Transfer) Final Report," Westinghouse Report WCAP-7665, April 1971). The use of a correlation derived from FLECHT data shall be demonstrated to be Refilling or reflooding is not required for the conservative for the transient to which it is applied; NuScale design as in a conventional PWR, presently available FLECHT heat transfer because there is no core uncovery (see the correlations ("PWR Full Length Emergency results of LOCA break spectrum calculations in Cooling Heat Transfer (FLECHT) Group I Test Section 9.0). This requirement is not relevant to Report," Westinghouse Report WCAP-7544, the NuScale LOCA EM. September 1970; "PWR FLECHT Final Report Supplement," Westinghouse Report WCAP-7931, Therefore, the required features of I.D.5.a are October 1972) are not acceptable. Westinghouse satisfied by design. Report WCAP-7665 has been approved for incorporation by reference by the Director of the Federal Register. A copy of this report is available for inspection at the NRC Library, 11545 Rockville Pike, Rockville, Maryland 20852-2738. New correlations or modifications to the FLECHT heat transfer correlations are acceptable only after they are demonstrated to be conservative, by comparison with FLECHT data, for a range of parameters consistent with the transient to which they are applied. I.D.5.b (Refill and Reflood Heat Transfer for Pressurized Water Reactors): Refilling or reflooding is not required for the NuScale design as in a conventional PWR, because there is no core uncovery (see the results During refill and during reflood when reflood rates of LOCA break spectrum calculations in Section are less than one inch per second, heat transfer 9.0). This requirement is not relevant to the calculations shall be based on the assumption that NuScale LOCA EM. cooling is only by steam, and shall take into account any flow blockage calculated to occur as a result of cladding swelling or rupture as such Therefore, the required features of D.5.b are blockage might affect both local steam flow and satisfied by design. heat transfer. The NuScale plant is not a BWR and does not I.D.6.Convective Heat Transfer Coefficients for have core spray cooling. Therefore, this Boiling Water Reactor Fuel Rods Under Spray requirement is not applicable to the NuScale Cooling. LOCA EM.

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10 CFR 50 Appendix K Required and NuScale LOCA EM Acceptable Feature The NuScale plant is not a BWR and does not I.D.7 The Boiling Water Reactor Channel Box have channel boxes. Therefore, this requirement Under Spray Cooling. is not applicable to the NuScale LOCA EM.

2.2.4 Other Requirements

Per the Design-Specific Review Standard for NuScale SMR Design, Section 4.4 (Reference 7), the thermal-hydraulic design should account for the effects of crud in the CHF calculations in the core or in the pressure drop throughout the RCS. NuScale will require that the fuel supplied for the NPM be supported with a qualified and approved product that supports this regulatory requirement. It is, however, acknowledged that crud deposition is driven by factors beyond fuel design, such as operating conditions and RCS chemistry. In order to evaluate the impact of crud on the LOCA FOMs, NuScale has evaluated the effect of the changes in thermal properties of the maximum credible crud thickness on fuel centerline and cladding temperatures during a LOCA. This evaluation determined that while the initial stored energy did increase as a result of crud, there was no significant impact on the LOCA response.

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3.0 NuScale Power Module Description and Operations

3.1 General Plant Design

The NuScale Power Plant consists of one or more Reactor Modules (RXM), each of which is a small, passive PWR. The RXM consists of the nuclear steam supply system (NSSS), which includes the nuclear core, the helical coil SGs and the pressurizer, within a single pressure vessel and the compact steel CNV that houses the NSSS.

Unique features of the NuScale plant design include the following:

• reduced core size • natural circulation reactor coolant flow (i.e., no reactor coolant pumps) • integrated SG and a pressurizer inside the RPV. As a result, there is no piping connecting the SG or pressurizer with the reactor • simplified passive safety-related systems that do not rely on ECCS pumps, accumulators, and water storage tanks (e.g., core makeup tank, in-containment refueling water storage tank) • high-pressure steel containment • containment immersed in a water-filled pool providing an effective passive heat sink for emergency cooling

The NPM is designed to operate efficiently at full-power conditions using natural circulation as the means of providing core coolant flow, eliminating the need for reactor coolant pumps. As shown in Figure 3-1, the reactor core is located inside a shroud connected to the hot leg riser. The reactor core heats reactor coolant, decreasing its density, causing the coolant to flow upward through the riser. When the heated reactor coolant exits the riser, it passes across the tubes of the helical coil SG, which acts as a heat sink. As the reactor coolant passes over the SG tubes, it cools, increases in density, and naturally circulates down the downcomer to the reactor core where the cycle begins again.

The NPMs are immersed in a reactor pool and protected by passive safety-related systems. Each NPM has a dedicated ECCS, chemical and volume control system (CVCS), and decay heat removal system (DHRS).

NuScale has achieved a substantial improvement in safety over existing plants through simplicity of design, reliance on passive safety-related systems, and small fuel inventory. The definition of a LOCA in 10 CFR 50.46(c)(1) addresses the geometry of a typical PWR, in which reactor coolant piping connects the RPV to primary system components external to the RPV. In the NuScale Power Plant design, all primary components are integral to the RPV, eliminating external coolant loops and pressurizer piping, which significantly reduces the number of possible LOCA scenarios.

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Figure 3-1. A single NuScale Power Module during normal operation

The potential break sizes included in the LOCA EM include the possible spectrum of breaks that can result in a break flow that exceeds the capability of the CVCS.

The NPM piping break locations are few (when compared to conventional PWR designs), and consist of the RCS injection and discharge lines, pressurizer spray supply line, and pressurizer high point vent line. These connections can be grouped into penetrations that are high on the RPV (pressurizer steam space) and low on the RPV (penetrate into an area which is normally in a liquid condition). All of the penetrations in the NPM design are at an elevation above the top of the core.

The NPM was designed with the intent of reducing the impact of a LOCA event. All LOCAs result in the actuation of both the ECCS and the DHRS. As shown in Figure 3-2, the ECCS consists of independent RVVs and independent RRVs.

The ECCS is initiated by opening the RVVs exiting the top of the RPV and the RRVs entering the RPV in the downcomer region (above the core elevation). Opening the valves allows the RPV and the CNV pressure to equalize which creates a natural circulation path to remove decay heat from the core. Water that is vaporized in the core leaves as steam through the RVVs, is condensed and collected in the CNV, and is then returned to the downcomer region inside the RPV through the RRVs by natural circulation.

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The CNV is sized such that the displacement of liquid from the RPV into containment will result in the liquid level being above the RRVs (which are located above the core) establishing a natural circulation loop. By the time the natural circulation pattern forms, the outside of the RPV will be cool enough that boiling on the outside of the RPV is relatively limited and the liquid level in the containment will have minimum swelling. The natural circulation loop removes decay heat from the core and RPV, and deposits it in containment. Heat deposited in containment is transferred by conduction and convection to the water in the reactor pool.

Following actuation of the ECCS, heat removal through the CNV rapidly reduces reactor and containment pressures and temperatures, and maintains them at acceptably low levels for extended periods of time. Because the CNV is evacuated to a low absolute pressure during normal operation (i.e., vacuum), only a small amount of non- condensable gas will be present inside the CNV at the beginning of the event.

The DHRS provides additional capacity to remove decay heat during the initial blowdown period of a LOCA, but it is neither required nor credited for such events. The DHRS provides secondary-side reactor cooling when normal feedwater is not available. The system, as shown in Figure 3-1, is a closed-loop, two-phase natural circulation cooling system. Two trains of decay heat removal equipment are provided, one attached to each SG loop. Each train is independently capable of removing 100 percent of the decay heat load and can cool the reactor primary-side inventory. Each train has a passive condenser submerged in the reactor pool. The condensers are maintained with sufficient water inventory for stable operation.

Analyses using the EM described in this report show that LOCAs do not challenge the safety of an NPM (see the results of LOCA break spectrum calculations in Section 9.0).

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Figure 3-2. Schematic of NuScale Power Module decay heat removal system and emergency core cooling system during operation

3.2 Plant Operation

This LOCA EM initiates the analyses of an NPM with 102 percent full-rated power operation (as required by 10 CFR 50 Appendix K). This assumption represents the uncertainty in the initial power.

Pressurizer heaters and a spray system are used to maintain nominal operating pressure similar to conventional PWRs. The reactor coolant is driven by natural circulation. At nominal full-power conditions, the flow rate is dependent on the fluid density differences through the loop, the losses incurred along the loop, and the elevation difference between the core and the SG.

During nominal full-power conditions, the control rods are retracted up to or above their insertion limits. Borated water is used as the primary coolant and the CVCS regulates the boron concentration to maintain criticality. The CVCS provides reactor inventory make-up through the RCS injection line in the riser and inventory let-down through a separate RCS discharge line in the downcomer region.

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The secondary side is operated such that the SGs remove the heat generated by the reactor core. The DHRS heat exchangers are isolated from the steam line and do not remove heat during normal operation.

The containment is evacuated during normal operation to provide an insulated barrier between the reactor and containment; no physical RPV insulation is present inside containment.

3.3 Safety-Related System Operation

The NuScale Module Protection System (MPS) is composed primarily of the reactor trip system and the engineered safety features actuation system. The MPS protection functions are limited to automated safety responses to off-normal conditions. The MPS functional response to an initiating event is a reactor trip; isolation (as necessary) of main feedwater, main steam, CVCS, and containment; followed by an integrated safety actuation of one or more of the passive safety-related systems (DHRS and ECCS). Containment isolation is achieved by closing of the following containment isolation valves:

• CVCS isolation valves - CVCS makeup line - CVCS letdown line - CVCS pressurizer spray supply line - CVCS high point degasification line, • reactor component cooling water system isolation valves • main steam system isolation valves • feedwater system isolation valves • containment flood and drain system isolation valves • containment evacuation system isolation valves

Dual safety-related isolation valves are installed on piping for the CVCS, containment evacuation system, containment flood and drain system, and reactor component cooling water system. There is one safety-related containment isolation valve in the main steam and feedwater piping penetrating containment with a redundant nonsafety-related isolation valve for each safety-related valve.

The reactor trip system consists of four independent separation groups with independent measurement channels to monitor plant parameters that can generate a reactor trip. Each measurement channel trips when the parameter exceeds a predetermined setpoint.

The engineered safety features actuation system also consists of four independent separation groups with independent measurement channels that monitor plant parameters that activate the operation of the engineered safety features.

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ECCS is actuated by MPS on high CNV level (interlocked with RCS hot temperature and pressurizer level) or low RCS pressure (interlocked with RCS hot temperature and CNV pressure). High CNV level or low RCS pressure are primary indications of a LOCA event. Interlocks are designed to prevent ECCS actuations for expected operational conditions or non-LOCA transients.

3.3.1 Emergency Core Cooling System

The ECCS is a two-phase natural circulation system that maintains a liquid water supply to the core during its operation in a LOCA scenario. This results in a collapsed liquid level in the RPV that is above the top of the core.

The ECCS consists of three independent RVVs and two independent RRVs. It is initiated by simultaneously actuating the RVVs on the top of the RPV in the pressurizer region and the RRVs on the side of the RPV in the downcomer region. The RRVs are designed to provide a low-resistance flow path for coolant to flow from the CNV into the RPV. The RVVs are designed to equalize pressure between the two vessels allowing steam from the reactor to vent to the containment and to provide hydrostatic equalization that allows coolant flow through the RRVs back into the reactor.

The ECCS actuation creates a steam flow path from the pressurizer to the containment and an RPV downcomer flow path to and from containment.

The RPV depressurizes due to liquid and steam exiting the ECCS valves. Steam entering containment is condensed on the containment wall, which in turn is cooled by the reactor pool. Initially, the containment pressure will increase to a peak, and then decrease as flow from the RPV decreases and heat is transferred from the CNV to the reactor pool. The RPV water inventory decreases while the containment level increases due to inventory transferred from the RPV.

As the pressure between the two vessels reach a near-equilibrium condition, the collapsed liquid level in the containment rises to a level higher than the RRV elevation, creating enough static head to overcome the pressure difference between the RPV and CNV. At this point, the condensed liquid in containment enters the RPV through the RRVs while steam exits the RPV through the RVVs. This stable process continues maintaining a collapsed water level above the top of the active fuel.

All ECCS valves are equipped with an inadvertent actuation block (IAB), the feature that prevents spurious opening of the ECCS valves at full operating pressure. The IAB prevents the valves from opening when the differential pressure between the RPV and CNV is greater than the IAB threshold pressure setpoint. After the IAB has blocked a spurious opening of the ECCS valve, it allows the valve to open only after the differential pressure between the RPV and CNV has decreased below the IAB release pressure setpoint.

The ECCS valves will also open on low differential pressure between the RPV and CNV, independent of an ECCS actuation signal. This action is a function of the mechanical design of the valves, where the valve spring will cause the valves to open if the pressure

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difference across the main chamber drops below approximately 15 psid. The spring force opening of the ECCS valves is expected during normal shutdown operations. The ECCS valves could open by this function during a rapid depressurization event where the RPV and CNV pressures equalize. This function could open ECCS valves earlier in some depressurization scenarios but does not significantly impact accident progression or results.

3.3.2 Decay Heat Removal System

The DHRS is a passive safety-related system that relies on natural circulation to remove heat from the RCS through the SG and reject heat to the reactor pool through the DHRS condenser. The DHRS is composed of two DHRS trains associated with one of the two NPM SGs. Each DHRS train is capable of independently removing 100 percent of decay heat. The DHRS piping connects to the main steam and feedwater lines specific to the associated SG. During normal operation, the DHRS condenser and piping are isolated by valves on the steam side of the SG. The condensate side of the DHRS is open to the feedwater piping supplying the associated SG.

Upon actuation of the DHRS, the SG feedwater and steam isolation valves close and the DHRS isolation valves open, creating a closed loop between the SG and DHRS condenser. Both liquid and vapor are contained in the DHRS on system actuation. Because the DHRS is a closed system, the total water mass remains constant during the system operation.

For successful operation, liquid water enters the SG through the feedwater line and is boiled by heat from the RCS. The vapor exits the SG through the steam line and is directed to the DHRS condenser where it condenses back to liquid before return to the SG. Thus, the loop transfers heat from the RCS to the DHRS fluid and then from the DHRS to the reactor pool water.

The bottom of the DHRS condenser is located above the bottom of the SG providing the static head to drive natural circulation.

The DHRS provides additional capacity to remove decay heat during the initial blowdown period of a LOCA. However, the break spectrum calculation (see Section 5.4) includes sampling conditions where both the DHRS trains are excluded. Not crediting DHRS operation provides results that cover the full range of possible DHRS performance conditions (including full failure).

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4.0 Phenomena Identification and Ranking

4.1 Phenomena Identification and Ranking Process

The purpose of the NuScale LOCA PIRT is to provide an assessment of the relative importance of phenomena and processes that may occur in the NPM during LOCA conditions in relation to specified FOMs. The PIRT assessment is part of the EMDAP process prescribed by RG 1.203.

The initial NuScale LOCA PIRT was developed in 2008. This PIRT was subsequently updated in 2013 and 2015 to address the design changes. The initial PIRT and the PIRT updates have been developed by a panel of recognized industry experts and NuScale subject matter experts, and are built upon the state-of-knowledge at the time of their development. The panel members of the initial LOCA PIRT were

• Dr. Brent Boyack (Los Alamos National Laboratory, retired) • Dr. Larry Hochreiter (Pennsylvania State University) • Dr. Mujid Kazimi (Massachusetts Institute of Technology) • Dr. Jose Reyes (NuScale Power, Inc.) • Dr. Kord Smith (Studsvik Scandpower, Inc.) • Dr. Graham Wallis, Chair (Darthmouth University, Creare, Inc.)

The panel members for the 2013 NuScale LOCA were

• Mr. Steve Congdon (GE Nuclear Energy, retired) • Dr. Tom George (Zachry Nuclear Engineering) • Mr. Craig Peterson (Computer Simulation and Analysis) • Dr. Jose Reyes (NuScale Power, Inc.) • Mr. Gregg Swindlehurst (GS Nuclear Consulting, LLC) • Dr. Graham Wallis, Chair (Darthmouth University, Creare, Inc.)

The panel members for the 2015 NuScale LOCA were

• Mr. Steve Congdon (GE Nuclear Energy, retired) • Dr. Tom George (Zachry Nuclear Engineering) • Dr. Jose Reyes (NuScale Power, Inc.) • Mr. Gregg Swindlehurst (Chair, GS Nuclear Consulting, LLC) • Dr. Graham Wallis, Chair (Darthmouth University, Creare, Inc.)

The 2015 NuScale LOCA PIRT incorporates lessons learned from testing and insights gained from computer code simulations of many LOCA scenarios.

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The PIRT panel received an in-depth briefing on the NPM design, LOCA sequence of events, and computer code predictions of the response of the NPM to LOCA scenarios. The panel then followed the PIRT process by first identifying the structures, systems, and components (SSC) of the NPM that were associated with the LOCA scenario. The LOCA scenario was then separated into phases with each phase representing a distinct process-dominated time period. Then FOMs were selected for each phase. Specifically, the FOMs were chosen to be quantifiable measures of the systems potential to meet regulatory safety limits. Phenomena were identified for each SSC for each phase, and the phenomena were ranked considering their level of importance relative to the FOMs. The panel also established a knowledge ranking for each of the phenomena.

The first NuScale LOCA PIRT was developed in 2008. This PIRT was subsequently updated in 2013 and 2015, primarily to address the changes in NPM design and operation. The PIRT panel was reconvened for each PIRT update and was presented with the changes in NPM design and their impact on progression of LOCA. The biographical information for each PIRT panel member is included with each PIRT release. The 2015 NuScale LOCA PIRT is used for the development of LOCA EM.

The following section provides a brief description of the LOCA scenarios and the accident phases considered for the PIRT developed. The definitions of the selected FOMs and the importance and knowledge ranking categories are summarized. Finally, the list of phenomena that were ranked as high importance by the PIRT panel in at least one of the phases of the NuScale LOCA scenarios is provided along with the brief description of the rational for assigned importance and knowledge level rankings. The rankings for all the identified phenomena and detailed description of the rationale are available in the 2015 NuScale LOCA PIRT report.

4.2 Loss-of-Coolant Accident Scenarios

Loss-of-coolant accidents are postulated breaks in the reactor coolant pressure boundary that result in leakage of reactor coolant at a rate exceeding the capability of the normal reactor coolant makeup system, as defined in 10 CFR 50.46(c)(1).

Breaks of various sizes, types, and orientations are postulated to occur in piping connected to the RPV. With the elimination of most primary coolant piping in the NPM design, breaks are limited to RCS injection and discharge lines, pressurizer spray supply line, and pressurizer high point vent line.

Two types of LOCA scenarios were addressed in the PIRT development process. The first type of LOCA scenarios were {{

}}2(a),(c) Section 9.1 provides further description of the progression of each LOCA scenario.

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The PIRT panel divided the NPM LOCA scenarios into two phases for the phenomena identification:

LOCA blowdown (Phase 1a)

Phase 1a begins with a postulated breach in the RCS pressure boundary that initiates the blowdown of the RCS into the CNV and ends when the MPS actuates ECCS to open the RVVs and RRVs.

ECCS actuation (Phase 1b)

Phase 1b begins when the MPS actuates ECCS to open the RVVs and RRVs and ends when the recirculation flow is established. The pressures and levels in containment and RPV approach a stable condition (i.e., initiation of long-term cooling).

Example calculations are performed for both steam space and liquid space LOCAs. Steam space breaks depressurize more quickly and generally actuate ECCS on low RCS pressure. Some liquid space breaks also actuate ECCS on low RCS pressure; however, the majority of the liquid space break spectrum actuates ECCS on high CNV level. The progression of the steam and liquid space LOCA events are similar, with the exception of different timing of the sequence of events and the liquid/steam composition of the break flow. The example calculations in this report were completed before the ECCS signal on low RCS pressure was added. The events that actuate earlier on low RCS pressure are less limiting with respect to the acceptance criteria. Therefore, the example calculations in this report are still representative and have conservative results.

4.3 Figures of Merit

The safe operation of the NPM was considered in the primary design phase. This produced a reactor system that protects the fuel using simple passive safety features. The NPM retains sufficient water in the RPV that the core will not be uncovered. For such a system, the LOCA PCT occurs at time zero (normal operating temperature). There is no heatup due to CHF or uncovering the core after event initiation; PCT remains below the 10 CFR 50.46 acceptance criterion of 2,200 degrees Fahrenheit (1,204 degrees Celsius) throughout the event. Hence, PCT is not an FOM for the NuScale PIRT process.

The critical heat flux ratio (CHFR) is an important FOM as it demonstrates there is no significant heatup of the cladding. One of the primary design fundamentals of the NPM is to protect the fuel from a CHF event. Therefore, an assessment of CHF becomes important.

Collapsed liquid level above the core is an additional FOM as it demonstrates there is an adequate supply of liquid water available to the core. Heatup of the fuel will not occur under LOCA conditions as long as the core is covered with coolant and CHF conditions do not exist.

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To ensure ECCS performance, the containment must be intact and remain below pressure and temperature design limits. Consequently, peak containment pressure and temperature are evaluated to ensure compliance with 50.46 criteria. However, the peak containment pressure and temperature for containment performance are calculated with a different methodology.

4.4 Definitions of Importance and Knowledge Level Rankings

Each phenomenon identified in the PIRT was assigned an importance ranking and knowledge level ranking. Table 4-1 and Table 4-2 describe the importance rankings and knowledge level rankings developed by the PIRT panel.

Table 4-1. Importance rankings

Importance Ranking Definition High (H) Significant influence on FOM Medium (M) Moderate influence on FOM Low (L) Small influence on FOM Inactive (I) Phenomenon not present or negligible

Table 4-2. Knowledge levels

Knowledge Level Definition 4 Well known/small uncertainty 3 Known/moderate uncertainty 2 Partially known/large uncertainty 1 Very limited knowledge/uncertainty cannot be characterized

4.5 Systems, Structures, and Components

To aid in the identification of phenomena, the PIRT panel divided the NPM into the SSC presented in Table 4-3. Phenomena were then identified in each SSC and for two of the LOCA phases.

Table 4-3. Systems, structures, and components

Reactor Pressure Decay Heat Removal Reactor Building Containment Vessel Vessel System Pool • • RPV heat source Reactor core • Decay heat removal Reactor Building pool • Containment vessel • Hot leg riser heat exchanger heat sink • DHRS isolation valves • Pressurizer • RRVs • SG • RRVs • RVVs • RVVs • Break • Break • Upper plenum • Downcomer

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Reactor Pressure Decay Heat Removal Reactor Building Containment Vessel Vessel System Pool • SG shell side (primary) • SG tube side (secondary) • LP

4.6 High-Ranked Phenomena

Separate PIRTs were developed for the two types of LOCA scenarios defined in Section 4.2. More than 80 phenomena were identified and ranked in each PIRT. Only a few differences were identified between the two LOCA scenarios with respect to the phenomena that might occur and their associated ranking. Table 4-4 summarizes the phenomena that were ranked as high importance by the PIRT panel in at least one of the two phases of the LOCA Scenarios 1 and 2. The knowledge level assigned by the PIRT panel is also included. These high-ranked phenomena are addressed in the development of NuScale LOCA EM. These phenomena and the rationale for their ranking are briefly described below.

Table 4-4. High-ranked phenomena

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4.6.1 Discussion of Phenomena Ranked High Importance

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4.7 Phenomena Identification and Ranking Table Summary

Some of the high-ranked phenomena identified in the NuScale PIRT are also important for existing reactors and have been the subject of considerable model development, testing, and analysis. Other phenomena are more unique to the NPM design due to the natural circulation coolant flow, integral RCS design, helical coil SG, unique passive ECCS and DHRS, reactor pool as the ultimate heat sink, and high-pressure containment. Phenomena associated with the helical coil SG and the DHRS are not ranked high importance because these systems do not play a significant role in determining the LOCA response with the assumptions used in this EM. {{

}}2(a),(c) However, the LOCA break analysis in Section 9.3 shows that not crediting the DHRS adds significant

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conservativism for these breaks. Since DHRS is assumed not available the phenomena are not ranked high importance.

Some of the unique phenomena have a more developed knowledge base due to occurrence of the phenomena in other designs with different geometries, e.g., natural circulation. The PIRT identified the phenomena within the specified components as the high-importance phenomena that have a low-knowledge level. These high importance, low knowledge phenomena are given the greatest focus in the development of the LOCA EM. {{

}}2(a),(c)

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5.0 Evaluation Model Description

This section provides a detailed description of the NPM LOCA model. The nodalization and modeling options selected for each NPM component are discussed along with the rationale for each choice. Justification is provided for the boundary and initial conditions selected for the model. A description of a break spectrum consistent with the requirements of 10 CFR 50 Appendix K is also provided. The NPM LOCA model is consistent with the SET and IET assessments used to validate NRELAP5 (see Section 7). The model follows the recommended best practices for the preparation of a RELAP5-3D© input (Reference 8) that are applicable to the NRELAP5 LOCA model, as well as the NuScale-specific LOCA guidelines summarized in this report. The model conforms to the applicable requirements of 10 CFR 50 Appendix K, as described in Section 2. The results of the break spectrum calculations and the sensitivity studies (i.e., nodalization, time step, initial and boundary conditions, and selected model parameters) that supported the development of the LOCA EM are summarized in Section 9. Specific initial and boundary condition values and the inputs for other key model parameters are specified in Appendix A.

5.1 NRELAP5 Loss-of-Coolant Accident Model for the NuScale Power Module

The unique design features of the NPM permit a simple and reliable approach to evaluate and mitigate the consequences of postulated LOCAs by:

• ensuring that all LOCAs are contained within the containment pressure vessel by designing the NPM such that the isolation of the CNV is a safety-related system. • actuating the ECCS valves, which depressurizes the RPV into the CNV to establish pressure equalization to allow return of discharged fluid back into the RPV to cool the core. • maintaining stable natural circulation flow through the ECCS valves with the reactor pool acting as the ultimate heat sink.

In the event of a LOCA, these design features result in a simple, predictable transient progression, that can be explained by a standard mass and energy balance over the RPV and CNV considering:

• choked or unchoked flow through the break and ECCS valves between RPV and CNV. • core decay heat generation and RCS stored energy release, heat transfer between CNV and reactor pool that is characterized by steam condensation at the CNV inside surface and free convection at the CNV outside surface to reactor pool.

5.1.1 General Model Nodalization

The NRELAP5 model for analyzing a NPM LOCA is developed by reviewing the postulated scenarios and the key phenomena described in the NuScale LOCA PIRT,

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summarized in Section 4. The model describes the key components of the NPM participating in a LOCA, as follows:

• RPV with internals - LP - reactor core - riser including the riser upper plenum - upper and lower downcomer - pressurizer • CNV • SG secondary side with DHRS condensers • reactor pool • ECCS valves • postulated break locations • RPV internal heat structures and heat structures between components (i.e., RPV to CNV to reactor pool) • riser holes - this feature location is approximately at the midpoint of the SG but is not included in the noding diagram. Evaluations determined riser holes do not significantly impact LOCA analysis results.

The nodalization diagram of these key components is shown in Figure 5-1. The details of the NRELAP5 NPM model are described in the following sections.

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{{

}}2(a),(c) Figure 5-1. Noding diagram of NRELAP5 loss-of-coolant accident input model for NuScale Power Module

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5.1.2 Reactor Coolant System

The RCS model is composed of the LP, reactor core, riser (lower, transition, and upper sections), riser plenum, downcomer (upper section containing the helical coil SGs and lower section), and pressurizer.

5.1.2.1 Lower Plenum

{{

}}2(a),(c)

5.1.2.2 Reactor Core

5.1.2.2.1 General Model

The reactor core assembly is modeled with {{

}}2(a),(c)

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Various passive heat structures inside the reactor core are also considered to increase the release of the stored energy accumulated in non-heat-generating structures with appreciable metal mass. These structures include:

• core support assembly including core barrel, reflector, upper support blocks, and lower core plate, • additional mass in fuel assemblies including top and bottom nozzles, upper and lower end caps, spacer grids, control rod assembly, instrument guide tubes, and springs.

5.1.2.2.2 Initial Power

In accordance with 10 CFR 50 Appendix K, Section I.A, the initial reactor power level is set to 102 percent of the rated thermal power to account for two percent measurement uncertainty. With a rated thermal power of 160 MWt, the initial reactor power before the initiation of a postulated LOCA is 163.2 MWt.

5.1.2.2.3 Core Power Distribution

The power distribution {{

}}2(a),(c) The sensitivity calculations presented in Section 9.6.6 show that axial power shape has negligible impact on LOCA FOMs.

5.1.2.2.4 Fuel Stored Energy

The fuel rods are initialized at the maximum initial stored energy condition as required by 10 CFR 50 Appendix K, Section I.A.1. The UO2 fuel thermal conductivity, volumetric heat capacity, and fuel-cladding gap conductance are considered to be the key thermo- physical properties defining the stored energy in the fuel, as the amount of stored energy is inversely proportional to the thermal diffusivity of the fuel and fuel-cladding gap conductance. Based on the burnup-dependent fuel performance analysis provided by the fuel vendor, choosing {{

}}2(a),(c) Additional details regarding the specification of thermal properties are provided in Appendix A.

All reactor core power is deposited into the fuel pellet directly, conservatively neglecting direct moderator heating and heat deposition to in-core materials. This approach is conservative as it maximizes the initial stored energy of the fuel.

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5.1.2.2.5 Point Kinetics Model and Decay Heat

The reactor kinetics model accounts for fission power due to prompt and delayed neutrons, decay power due to fission products, and actinides.

The reactivity equations are solved using a point kinetics model with the ‘separable’ option (see Section 6.4) to calculate the fission power. This model simulates reactivity changes due to reactor trip and feedback reactivity due to Doppler and moderator density effects. The reactivity change due to the insertion of control rods is modeled using a scram reactivity table as a function of time after the reactor trip. The table reflects conservative representations for the onset of rod motion, the rod position as a function of time after trip and the inserted reactivity as a function of rod position. The scram rod total worth considers that the most reactive rod remains stuck and does not insert into the core following reactor trip. Sensing signal delays are accounted for based on the instrument type (e.g., pressure and temperature) to determine if a reactor trip should be initiated and an additional delay is added to account for the initiation of rod insertion.

The Doppler and moderator density feedback input parameters depend on the average fuel burn-up. The weighted average fuel temperature and moderator density for the feedback reactivity calculations consider the given core power (flux) distribution to be consistent with the point kinetics model. {{

}}2(a),(c)

Six groups of delayed neutron precursors are considered based on the SIMULATE analysis for different cycles, beginning-of-cycle, middle-of-cycle, and end-of-cycle. Beginning-of-cycle kinetic parameters are used as bounding for point kinetic input with the smallest prompt neutron lifetime to maximize initial energy inventory by prolonging the fission power transient. The variation in precursor decay constants is insensitive to the time in cycle. Additional biasing is introduced to all kinetic parameters to account for uncertainty in calculated values in such a way that both the prompt lifetime and the decay constants for each precursor group are decreased. The objective is to maximize the delayed neutron contribution to the total fission power.

The ‘gamma-ac’ option in the NRELAP5 point kinetics model activates the models for the transient effects of decay heat and actinides. Use of the actinides model complies with 10 CFR 50 Appendix K, Section I.A.3. The ANS71 option represents an explicit implementation of four time-dependent exponentials detailed in the draft ANS 73 decay heat standard and includes a built-in 1.2 multiplier. Activation of a trip associated with this option initiates evaluation of the power decay as a function of time. However, this time-evaluation of power does not account for post-trip prompt and delayed fissions that can add additional decay heat precursors and increase the integral heat release.

The ANS73 model represents an 11 group exponential fit of production and decay constants of the decay heat defined in the standard. User input of a 1.2 multiplier addresses 10 CFR 50 Appendix K requirements for addressing prediction uncertainties (10 CFR 50 Appendix K, Section I.A.4). Decay heat is predicted by the behavior of the

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11 precursor groups, and no explicit reactor trip is applied. Instead the model predicts precursor concentrations from the prompt and delayed fissions rate, and so naturally follows fission power.

Sensitivities have determined that the ANS71 option requires careful selection of an additional delay time beyond a reactor trip before activating the power decay to account for post-trip prompt and delayed fissions. This delay time, which depends on control rod insertion speed and reactivity, can be more than {{ }}2(a),(c) seconds. Once the delay is accounted for, both options are consistent for short term LOCA through the first 1000 seconds. Considering the correct response of the ANS73 option without any special evaluation of delay times, the ANS73 option was chosen as the standard choice in the LOCA guideline.

The best-estimate ANS79 actinides model is used to account for heat deposition from actinide decay. The actinide model includes the decay energies from the production and decay of 239U, 239Np, and 239Pu.

5.1.2.3 Riser

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5.1.2.4 Downcomer

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5.1.2.5 Pressurizer

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5.1.2.6 Collapsed Liquid Level Calculation

The collapsed liquid levels in the riser, downcomer, and containment are calculated based on the total liquid volume calculated in each part of the NPM and volume-elevation table for these regions. The riser volume includes the lower plenum, core and bypass region, and riser section up to the pressurizer baffle plate1. Similarly, the downcomer volume includes the lower plenum, lower and upper downcomer section, and upper riser plenum up to the bottom of the pressurizer or the pressurizer baffle plate1. This approach to collapsed liquid level will allow for there to be substantial flashing and momentary voiding in the core, such as that seen at near stagnant flow conditions for small liquid space breaks of less than 35 percent. This is discussed in

1 The riser and downcomer collapsed liquid level calculations in the final FSAR analysis do not include the upper plenum volume (node 140-1 in Figure 5-1).

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Section 9.2 where assurance of no fuel CHF, and hence no fuel heat-up, is shown with transient MCHFR remaining above the steady-state CHFR value for all cases.

5.1.3 Helical Coil Steam Generators

Two helical coil SGs are represented using an NRELAP5-specific helical SG component that models the component-specific internal pressure drop and heat transfer effects, as described in Section 6.7. The two independent SGs are thermally connected to the upper downcomer to transfer heat to the steam turbine during normal operation. During off-normal operations, each SG transfers energy to an independent safety-related DHRS (see Section 5.1.7) to discharge energy to the reactor pool.

The tube-to-coil diameter ratio is specific to the SG geometry. {{

}}2(a),(c)

5.1.4 Containment Vessel and Reactor Pool

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The reactor pool represents the ultimate heat sink in the NuScale design. The reactor pool volume corresponding to an individual NPM is represented by a {{

}}2(a),(c) A wide range of initial reactor pool temperatures is exercised to show the effect of the pool conditions on the LOCA behavior in Section 9.6.5.

{{

}}2(a),(c)

5.1.5 Chemical and Volume Control System

The entirety of the CVCS is not explicitly included in the LOCA model. The CVCS, a nonsafety-related system, is not automatically actuated. {{

}}2(a),(c)

Continued operation of the CVCS through operator action would add cold water that is non-conservative. The only CVCS piping represented in the model is the injection line from the RPV wall, through the downcomer and into the riser. It connects the charging line break to the containment vessel at the correct elevation and accounts for a small loss through the line. The discharging line connection at the downcomer, and the two spray supply line and high point vent line connections at the top of the pressurizer are used as break locations with no attached piping included. The volumes corresponding to removed CVCS piping constitute a small fraction of the total RPV and CNV volume; therefore, it has negligible impact on the progression of NPM LOCA.

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5.1.6 Secondary System

The model represents the secondary feedwater and steam lines with two helical coil SGs, described in Section 5.1.3. {{ }}2(a),(c) The secondary side includes the DHRS with two trains of heat exchangers with feed and steam line piping, described in Section 5.1.7.

{{

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5.1.7 Decay Heat Removal System

Both DHRS trains are included in the NPM LOCA model. The two independent trains of the DHRS are safety-related systems; however, no credit is taken for operation of the DHRS in the LOCA methodology. The break spectrum calculation results discussed in Section 9.3 confirm that this assumption is highly conservative for LOCA analysis.

5.1.8 NRELAP5 Modeling Options

The NPM LOCA analysis is performed with the latest released version of NRELAP5. {{

}}2(a),(c)

5.1.8.1 Junction Options

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Table 5-1. Default junction options for the NRELAP5 loss-of-coolant accident model

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5.1.8.2 Volume Options

{{ }}2(a),(c) This format is described by Table 5-2.

Table 5-2. Default volume options for the NRELAP5 loss-of-coolant accident model

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5.1.8.3 Heat Structure Options

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5.1.9 Time Step Size Control

The NuScale LOCA EM uses the NRELAP5 semi-implicit scheme for the solution of the hydrodynamics. The heat structure solution is implicitly coupled to the hydrodynamic solution. With given user-specified minimum and maximum time step sizes, the code determines the appropriate time step in such a way that

• the current time step cannot be larger than the courant-time step size determined based on the limiting volume.

• no significant mass error accumulation occurs during the solution and halving of the current time step when it is deemed necessary.

NRELAP5 provides the capability of providing the user-defined maximum time step size through the definition of control variable. A control variable that defines the fraction of the current courant time-step size during the solution is used to set the user-defined maximum time step size. This approach has the advantage of taking larger time steps when larger courant time step sizes exist during the solution; therefore, the code takes larger time steps when the solution indicates smooth transient progression. {{

}}2(a),(c)

A sensitivity study is performed on the fraction specified to demonstrate that the selected maximum time-step size has no or insignificant impact on the LOCA figures of merit such as peak containment pressure and collapsed liquid level in the RPV riser above the TAF.

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5.2 Analysis Setpoints and Trips

A number of safety-related measurements exist in the NPM to detect off-normal conditions. Table 5-3 shows the measurements relevant to LOCA analysis along with their functions.

Table 5-4 presents the list of actuation signals for the NPM safety-related systems and identifies the signals that are credited and not credited in the LOCA EM. The table footnotes provide important definitions for containment isolation and DHRS actuation. {{

}}2(a),(c) Not crediting these and other measurements for the associated safety-related signals would delay the activation of reactor trip and the DHRS under certain conditions, which is conservative for LOCA analysis.

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Table 5-3. NuScale Power Module safety-related system measurement parameters

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Table 5-4. Safety-related system actuation signals

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{{

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The safety analysis analytical limits specify the setpoints (or range of setpoints) and the sensing delay for each safety-related signal. Table A-3 shows the setpoint values or analytical limits and the signal actuation delays used for the LOCA break spectrum calculation and the sensitivity calculations presented in Section 9.0. Table 5-5 shows the basis for the selection of the safety-related signal delays in the NuScale LOCA EM. Signals not credited either do not play a role in LOCA {{ }}2(a),(c) or act to provide additional conservatism in the delay of actuating safety-related systems that are only beneficial to the LOCA progression.

Table 5-5. Safety-related analysis signal delays {{

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The mixture level detection uses a simple approximation of the mixture level based on {{

}}2(a),(c)

5.3 Initial Plant Conditions

Table 5-6 provides the basis for conservatively biasing the initial conditions for LOCA analysis. Table A-2 of Appendix A provides the specific ranges of NPM primary and secondary side initial or operational conditions. These ranges are intended to account for both the normal control system deadband and the system/sensor measurement uncertainty without specifically quantifying the portion of the range applied to either uncertainty.

Table 5-6. Plant initial conditions

{{

}}2(a),(c) 5.4 Loss-of-Coolant Accident Break Spectrum

10 CFR 50 Appendix K describes the break spectrum as a set of LOCA scenarios that are uniquely defined based on location, configuration and size. Additional sensitivity studies were performed on availability of DHRS, availability of power, and postulated

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single failures. The break spectrum for the NuScale LOCA EM is summarized in this section.

5.4.1 Break Location

The postulated break locations in the NPM design are the RCS injection and discharge lines, the pressurizer spray supply line, and high point vent lines. These break locations establish a flow path between RPV and CNV leading to CNV pressurization during the early phase of LOCA (i.e., Phase 1a):

• The injection line enters the RPV through a shell penetration and piping internal to the RPV that passes through the downcomer and terminates at an upwardly-oriented nozzle in the riser. • The discharge line is connected to a RPV penetration to the downcomer. • The pressurizer spray supply line is connected to the top of the pressurizer at two separate penetrations. At each penetration there is a nozzle within the RPV wall. Outside the RPV wall (but within the CNV), the pressurizer spray supply lines connect to a tee which in turn connects to isolation valves on the CNV wall. • The high point vent line connects to the top of the pressurizer.

All of the connections to the RPV are normally open, except for the high point vent. Each connection can be isolated by two independent safety-related isolation valves in series that close on the containment isolation signal. As a result, all discharged break fluid is retained within the CNV for eventual return to the RPV when ECCS actuates.

5.4.2 Break Configuration and Size

Table 5-7 summarizes the size and location of the breaks considered as part of the break spectrum of the NuScale LOCA EM. {{

}}2(a),(c),ECI

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{{

}}2(a),(c),ECI

Table 5-7. Summary of analyzed break sizes {{

}}2(a),(c),ECI 5.4.3 Single Failures

10 CFR 50 Appendix K requires that single failures be considered within the break spectrum. This includes analyzing a system/component classified as nonsafety related if the inclusion of that system/component would introduce a more limiting condition. The following scenarios are considered:

• no single failure • failure of a single RVV to open • failure of a single RRV to open • failure of one ECCS division (i.e., one RVV and one RRV)

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The ECCS valves are held closed with direct current (DC) power and operate on two independent divisions. Each division controls one RRV and one RVV. The third RVV is connected to both Division 1 and 2. Removal of DC power from any division will cause the solenoid to release for this third RVV (ready to open pending dropping below the IAB release pressure). Two modes of failure can be postulated: (1) failure to actuate a division upon actuation request and (2) inadvertent actuation of a division.

In the event of failure of the actuation of a division, the DC power from one division is not removed when it should be removed. The result is one RRV and two RVVs actuating, with one RRV and one RVV not actuating. This scenario is explicitly analyzed within the LOCA EM.

In the event of an inadvertent actuation of a division (removal of DC power from that division) two RVVs and one RRV will be available to open immediately. The IAB setpoint will prevent the opening of these valves until the differential pressure between RCS and CNV falls below the IAB release pressure. If DC power is not available, all ECCS valves will open at the IAB release pressure. This case is covered as discussed in Section 5.4.4. If DC power is available the other division will still actuate on the level signal creating a staggered release. This is a non-limiting case as a staggered release has a smaller impact on system pressures, levels, and core coolability relative to the immediate opening of all ECCS valves at the IAB release pressure.

5.4.4 Loss of Power

Coincident with a postulated LOCA, two scenarios for loss of power are considered within the LOCA methodology:

• complete loss of normal alternating current (AC) and DC power and • complete loss of only AC power with DC power availability.

The loss of DC power can impact the LOCA progression by immediately triggering valves to go to their fail-safe position. Table 5-8 presents the valves along with their fail- safe state.

Table 5-8. NuScale Power Module valve fail-safe positions with loss of DC power {{

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For the ECCS valves after loss of DC power, the IAB arming valves close because the valve differential pressure is greater than the threshold setpoint, thereby preventing the immediate opening of the ECCS valves. As the RPV pressure decreases and the CNV pressure increases, the valves open as soon as the differential pressure drops below the IAB release pressure setpoint.

When normal AC power is lost, the feedwater pumps coast down and the turbine trip is initiated. Upon loss of normal AC power (with a time delay) the reactor trip, containment isolation, and DHRS actuation signals are generated (see Table 5-4 and Table 5-5). The ECCS does not receive an actuation signal based on loss of AC power until the 24 hour time delay is surpassed. The ECCS is still available to actuate based on its normal actuation signals (Table 5-4).

5.4.5 Decay Heat Removal System Availability

When the SG tubes are uncovered within the RPV, operation of the DHRS results in condensation of steam on the external surface of the helical coil SGs and retains liquid inventory in the RPV instead of releasing it to the CNV through vaporization. The RPV pressure is reduced for cases with DHRS available and there is a higher minimum inventory in the RPV.

There is no single failure that can prevent a single DHRS train from actuating; however, to account for uncertainties in modeling of the DHRS, the DHRS performance is considered in a sensitivity study. Specifically, consideration is given to full availability and full loss of both DHRS trains and is included in the LOCA break spectrum calculations presented in Section 9.3. The crediting of DHRS is not required for the LOCA EM to meet acceptance criteria.

5.5 Sensitivity Studies

The sensitivity calculations described in Section 9.6 are performed in three categories:

• sensitivies required by 10 CFR 50 Appendix K, ( e.g., nodalization and time-step size to demonstrate the stability and consistency of the numerical scheme used by the NRELAP5 code), • sensitivities related to key phenomena and design input parameters considered to be important to the LOCA progression and LOCA FOMs (e.g., CCFL at pressurizer baffle plate, ECCS parameters, etc.). • sensitivities to determine input parameters to ensure conservativism (e.g., reactor pool initial temperature, core power distribution, etc.).

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6.0 NRELAP5 Code Description

The NuScale LOCA EM is based on the NRELAP5 system thermal-hydraulics code. The NRELAP5 code includes models for characterization of hydrodynamics, heat transfer between structures and fluids, modeling of fuel, reactor kinetics models, and control systems. NRELAP5 uses a two-fluid, non-equilibrium, non-homogenous model to simulate system thermal-hydraulic responses. This section provides a general overview of the code structure, models, and correlations. This section also addresses the LOCA- specific code models and improvements implemented to address unique design features and phenomena for the NPM. The adequacy of code models and correlations essential for modeling all high-ranked PIRT phenomena is discussed in Section 8.0. The full details of the models and correlations that makeup NRELAP5 can be found in the NRELAP5 Theory Manual (Reference 9).

RELAP5-3D©, version 4.1.3, was used as the baseline development platform for the NRELAP5 code. RELAP5-3D© was procured and as part of the procurement process commercial grade dedication was performed by NuScale to establish the baseline NRELAP5 code. Subsequently, features were added and changes made to NRELAP5 to address the unique aspects of the NPM design and licensing methodology. Those aspects of NRELAP5 that are new or revised specifically for the NPM application include:

• {{

}}2(a),(c)

The previous RELAP5 series of codes were developed at the INL under sponsorship of the DOE, the U.S. NRC, members of the International Code Assessment and Applications Program, members of the Code Applications and Maintenance Program, and members of the International RELAP5 Users Group. Specific applications of the code have included simulations of transients in light water reactor systems, such as LOCAs, anticipated transients without scram, and anticipated operational occurrences, such as loss of feedwater, loss of offsite power, station blackout, and turbine trip.

The RELAP5 code, including the RELAP5-3D© version that was used as the development platform for NRELAP5, has an extensive record of usage and acceptable performance for nuclear safety analysis. RELAP5-3D© is the latest version of the RELAP5 code that has been under continuous development since 1975, first under NRC sponsorship and then with additional DOE sponsorship beginning in the early 1980s. While NRC sponsorship ended in 1997, the DOE continued sponsorship of RELAP5-3D© to meet its own reactor safety assessment needs. The RELAP5 code was chosen by DOE as the thermal-hydraulic analysis tool because of its widespread acceptance.

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Systematic safety analyses were carried out for the DOE that included the N reactor at Hanford, the K and L reactors at Savannah River, the Advanced Test Reactor at the Idaho National Engineering and Environmental Laboratory, the High Flux Isotope Reactor and Advanced Neutron Source at Oak Ridge, and the High Flux Beam Reactor at Brookhaven. The DOE also chose RELAP5 for the independent safety analysis of the New Production Reactor proposed for Savannah River .

RELAP5-3D© has worldwide usage for nuclear safety analysis. Users participate in the International RELAP5 Users Group (IRUG) which provides a forum for code users to share their RELAP5 development and analysis experiences. Meeting participants also communicate new features and applications that have been developed for RELAP5-3D©. Code users include reactor vendors, nuclear industry suppliers, a naval nuclear propulsion laboratory, universities, and international organizations. NuScale is a participant in IRUG.

RELAP5-3D© has been chosen as a code development platform for small break LOCA analysis by Mitsubishi Heavy Industries for APWR (Reference 21). Furthermore, U.S. NRC has performed a detailed adequacy evaluation of RELAP5/MOD3 Version 3.2.1.2 for analysis of design-basis small break LOCA in the Westinghouse AP600 reactor (Reference 74). This usage of RELAP5-3D© over a long period of time has produced a large amount of user feedback. Submission of code error reports and the follow up code development has resulted in a robust code which can be used with a high level of confidence that significant code problems have been identified and corrected.

The more than 18 year history of code assessment and successful application of the RELAP5-3D© code, and codes based on the RELAP5-3D© platform, by the worldwide user community has successfully exercised the fundamental capabilities of RELAP5-3D© that are the critical characteristics required of NRELAP5 for NuScale’s application.

The NRELAP5 code is developed following the requirements of the NuScale QAPD (Reference 4). The NuScale corporate Software Configuration Management Plan provides a framework for NRELAP5 configuration management and change control in conformance with the requirements outlined in the NuScale Software Program Plan. Review and approval of the NuScale corporate Software Configuration Management Plan is not within the scope of this report.

6.1 Quality Assurance Requirements

The NuScale QAPD complies with the requirements of 10 CFR 50 Appendix B, Quality Assurance Criteria for Nuclear Power Plants and Fuel Reprocessing Plants (Reference 10) and American Society of Mechanical Engineers (ASME) NQA-1-2008 and NQA-1a- 2009 Addenda, “Quality Assurance Program Requirements for Nuclear Facility Applications,” (Reference 12).

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6.2 NRELAP5 Hydrodynamic Model

The NRELAP5 hydrodynamic model is a transient, two-fluid model for flow of a two phase vapor/gas-liquid mixture that can contain non-condensable components in the vapor/gas phase as well as a soluble component (i.e., boron) in the liquid phase.

The two-fluid equations of motion that are used as the basis for the NRELAP5 hydrodynamic model are formulated in terms of volume and time-averaged parameters of the flow. Phenomena that depend upon transverse gradients, such as friction and heat transfer, are formulated in terms of the bulk properties using empirical transfer coefficient formulations. In situations where transverse gradients cannot be represented within the framework of empirical transfer coefficients, such as subcooled boiling, additional models specially developed for the particular situation are employed. The system model is solved numerically using a semi-implicit, finite-difference technique.

6.2.1 Field Equations

The NRELAP5 thermal-hydraulic model solves eight field equations for eight primary dependent variables. The primary dependent variables are pressure, phase-specific internal energies, vapor or gas volume fraction, phasic velocities, non-condensable quality, and boron density. For the one-dimensional equations, the independent variables are time and distance. Non-condensable quality is defined as the ratio of the non- condensable gas mass to the total vapor or gas phase mass.

The secondary dependent variables used in the equations are phasic densities, phasic temperatures, saturation temperature, and non-condensable mass fraction in the non- condensable gas phase for the ith non-condensable species.

The basic field equations for the two-fluid, non-equilibrium model consist of two phasic continuity equations, two phasic momentum equations, and two phasic energy equations. The equations are time averaged and one-dimensional. The phasic continuity equations are shown in Equation 6-1 and Equation 6-2.

𝜕 1 𝜕 𝛼 𝜌 + 𝛼 𝜌 v 𝐴 = Γ Equation 6-1 𝜕𝑡 𝐴 𝜕𝑥

𝜕 1 𝜕 𝛼 𝜌 + 𝛼 𝜌 v 𝐴 = Γ Equation 6-2 𝜕𝑡 𝐴 𝜕𝑥

Continuity consideration yields the interfacial condition of Equation 6-3.

Equation 6-3 Γ =−Γ

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The interfacial mass transfer model assumes that total mass transfer can be partitioned into mass transfer at the vapor/liquid interface in the bulk fluid (Γ) and mass transfer at the vapor/liquid interface in the thermal boundary layer near the walls (Γ) as defined by Equation 6-4.

Γ =Γ +Γ Equation 6-4

The phasic momentum equations are in the form of Equation 6-5 and Equation 6-6.

𝜕v 1 𝜕v 𝜕𝑃 𝛼 𝜌 𝐴 + 𝛼 𝜌 𝐴 =−𝛼 𝐴 +𝛼 𝜌 𝐵 𝐴−𝛼 𝜌 𝐴𝐹𝑊𝐺 ∙ v 𝜕𝑡 2 𝜕𝑥 𝜕𝑥

+Γ𝐴v − v−𝛼𝜌𝐴𝐹𝐼𝐺 ∙v − v Equation 6-5

𝜕v − v 𝜕v 𝜕v −𝐶𝛼 𝛼 𝜌 𝐴 + v − v 𝜕𝑡 𝜕𝑥 𝜕𝑥

v v 𝛼 𝜌 𝐴 + 𝛼 𝜌 𝐴 =−𝛼 𝐴 +𝛼 𝜌 𝐵 𝐴−𝛼 𝜌 𝐴𝐹𝑊𝐹 ∙ v

−Γ𝐴v − v−𝛼𝜌𝐴𝐹𝐼𝐹 ∙v − v Equation 6-6

𝜕v − v 𝜕v 𝜕v −𝐶𝛼 𝛼 𝜌 𝐴 + v −𝑣 𝜕𝑡 𝜕𝑥 𝜕𝑥

The force terms on the right sides of Equation 6-5 and Equation 6-6 are, respectively, the pressure gradient, the body force (i.e., gravity and pump head), wall friction, momentum transfer due to interface mass transfer, interface frictional drag, and force due to virtual mass. The terms 𝐹𝑊𝐺 and 𝐹𝑊𝐹 are part of the wall frictional drag, which are linear in velocity, and are products of the friction coefficient, the frictional reference area per unit volume, and the magnitude of the fluid bulk velocity. The coefficients 𝐹𝐼𝐺 and 𝐹𝐼𝐹 are part of the interface frictional drag; two different models (drift flux and drag coefficient) are used for the interface friction drag, depending on the flow regime.

Conservation of momentum at the interface requires that the force terms associated with interface mass and momentum exchange sum to zero as shown by Equation 6-7.

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𝜕v − v 𝛤 𝐴𝑣 𝛼 𝜌 𝐴𝐹𝐼𝐺v ∙ − v −𝐶𝛼 𝛼 𝜌 𝐴 𝜕𝑡 Equation 6-7

vv −𝛤 𝐴𝑣 𝛼 𝜌 𝐴𝐹𝐼𝐹 ∙v − v −𝐶𝛼 𝛼 𝜌 𝐴 =0

The phasic thermal energy equations are defined by the following two equations:

𝜕 1 𝜕 𝜕𝛼 P 𝜕 𝛼 𝜌 𝑈 + 𝛼 𝜌 𝑈 v 𝐴 = −P − 𝛼 v A 𝜕𝑡 𝐴 𝜕𝑥 𝜕t A 𝜕x Equation 6-8

∗ ` +Q +𝑄 − Γh − Γh +𝐷𝐼𝑆𝑆

∂ 1 ∂ ∂α P ∂ α ρ U + α ρ U v A = −P − (α v A) ∂t A ∂x ∂t A ∂x Equation 6-9

∗ ` +Q +Q − Γh − Γh +DISS

In the phasic energy equations, Q and Q are the phasic wall heat transfer rates per unit volume. These phasic wall heat transfer rates satisfy Equation 6-10 where Q is the total wall heat transfer rate to the fluid per unit volume.

𝑄=𝑄 +𝑄 Equation 6-10

The vapor generation (or condensation) consists of two parts, vapor generation that results from energy exchange in the bulk fluid (Γ) and energy exchange in the thermal boundary layer near the wall (Γ) (Equation 6-4). Each of the vapor generation (or condensation) processes involves interface heat transfer effects. The interface heat transfer terms (𝑄 and Q) appearing in Equation 6-8 and Equation 6-9 include heat transfer from the fluid states to the interface due to interface energy exchange in the bulk and in the thermal boundary layer near the wall. The vapor generation (or condensation) rates are established from energy balance considerations at the interface.

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The phasic energy dissipation terms, DISS and DISS, are the sums of wall friction, pump, and turbine effects. The dissipation effects due to interface mass transfer, interface friction, and virtual mass are neglected.

6.2.2 State Relations

The six-equation model uses five independent state variables with an additional equation for the non-condensable gas component. The independent state variables are chosen to be 𝑃, 𝛼, 𝑈, 𝑈, and 𝑋. All the remaining thermodynamic fluid variables (temperatures, densities, partial pressures, qualities, etc.) are expressed as functions of these five independent state variables. In addition to these variables, several state derivatives are needed for some of the linearizations used in the numerical scheme.

𝜕𝜌𝑔 𝜕𝜌𝑔 𝜕𝜌𝑔 𝜕𝜌𝑓 𝜕𝜌𝑓 Equation 6-11 , , , , 𝜕𝑃 𝜕𝑈𝑔 𝜕𝑋𝑛 𝜕𝑃 𝜕𝑈𝑓 𝑈𝑔,𝑋𝑛 𝑃,𝑋𝑛 𝑃,𝑈𝑔 𝑈𝑓 𝑃

The interphase mass and heat transfer models use an implicit (linearized) evaluation of the temperature potentials 𝑇 −𝑇 and 𝑇 −𝑇. The quantity 𝑇 is the temperature that exists at the phase interface. The implicit (linearized) evaluation of the temperature potentials in the numerical scheme requires the derivatives of the phasic and interface temperatures defined by Equation 6-12.

𝜕𝑇 𝜕𝑇 𝜕𝑇 𝜕𝑇 𝜕𝑇 𝜕𝑇 𝜕𝑇 𝜕𝑇 Equation 6-12 , , , , , , , 𝜕𝑃 𝜕𝑈 𝜕𝑋 𝜕𝑃 𝜕𝑈 𝜕𝑃 , 𝜕𝑈 𝜕𝑋 , , , , ,

6.2.2.1 Water Property Tables

The set of basic properties for light water is used for all LOCA calculations. Implementation is activated by the user. These thermodynamic tables tabulate saturation properties as a function of temperature, saturation properties as a function of pressure, and single-phase properties as a function of pressure and temperature. The tables are based on the 1995 Steam Tables from the International Association for the Properties of Water and Steam (IAPWS) and are known as IAPWS-95. The temperature and pressure range covered in the property table is 273.16 K (32.018 degrees F) to 5000 K (8540.33 degrees F) and 611.6 Pa (0.0887 psia) to 100 MPa (14,504 psia). The properties and derivatives in the tables are saturation pressure, saturation temperature, specific volume (υ), specific internal energy, specific entropy, and three derivatives: the isobaric thermal expansion coefficient (β), the isothermal compressibility (κ), and the specific heat at constant pressure (Cp).

6.2.2.2 Single-Component, Two-Phase Mixture

Liquid properties are obtained from the thermodynamic tables, given P and Uf. All the desired density and temperature derivatives can then be obtained from the derivatives of

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κf, βf, and Cpf . In the case of the vapor being subcooled or the liquid being superheated, (i.e., metastable states) the calculation of υ, T, κ, β, and Cp incorporates a constant pressure extrapolation from the saturation state for the temperature and specific volume.

6.2.3 Flow Regime Maps

The one-dimensional nature of the field equations for the two-fluid model used in NRELAP5 precludes direct simulation of effects that depend upon transverse gradients of any physical parameter, such as velocity or energy. Consequently, such effects must be accounted for through algebraic terms added to the conservation equations.

The mapping for flow conditions to a specific flow regime is required to provide closure to the two-fluid equations. The selected flow regime determines the constitutive relationships that are applied for interphase friction, the coefficient of virtual mass, wall friction, wall heat transfer, and interphase heat and mass transfer. The flow regime maps are based on the work of Taitel and Dukler (References 14 and 15) and Ishii (References 16, 17, and 18). Taitel and Dukler have simplified flow regime classifications and developed semi-empirical relations to describe flow regime transitions. However, some of their transition criteria are complex, and further simplification has been carried out in order to efficiently apply these criteria in NRELAP5.

The flow regime maps for the volumes and junctions are identical but used differently as a result of the finite difference scheme and staggered mesh used in the numerical scheme. The volume map is based on volume quantities. It is used for interphase heat and mass transfer, wall friction, and wall heat transfer. Meanwhile, the junction map is based on junction quantities and is used to calculate the interfacial friction coefficient.

Three flow-regime maps in both volumes and junctions for two-phase flow are used in the NRELAP5 code: (a) a horizontal map for flow in pipes; (b) a vertical map for flow in pipes, annuli, and bundles; and (c) a high mixing map for flow through pumps.

Wall heat transfer depends on the volume flow regime maps in a less direct way. Generally, void fraction and mass flux are used to incorporate the effects of the flow regime. Since the wall heat transfer is calculated before the hydrodynamics, the flow information is taken from the previous time step.

6.2.3.1 Vertical Volume Flow Regime Maps

The vertical volume flow regime map is for upflow, downflow, and counter current flow in volumes whose inclination (vertical) angle 𝜙 is such that 60 < |𝜙| ≤90 degrees. An interpolation region between vertical and horizontal flow regimes is used for volumes whose absolute value of the inclination (vertical) angle is between 30 and 60 degrees.

This map is modeled as nine regimes:

• four regimes for pre-CHF heat transfer - bubbly, slug, annular-mist, and dispersed (droplet or mist)

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• four regimes for post-CHF heat transfer - inverted annular, inverted slug, mist, and dispersed (droplet or mist) • one regime for vertical stratification

A schematic of the vertical flow regime map as coded in NRELAP5 is shown in Figure 6-1. The schematic is three-dimensional to illustrate flow-regime transitions as functions of void fraction (𝛼), average mixture velocity (𝑣), and boiling.

Figure 6-1. Schematic of vertical flow-regime map indicating transitions

6.2.3.2 Junction Flow Regime Maps

The junction map is based on both junction and volume quantities. It is used for the interphase drag and shear, as well as the coefficient of virtual mass. The flow regime maps used for junctions are the same as used for the volumes and are based on the work of Taitel and Dukler (Reference 14 and Reference 15), Ishii (Reference 16), and Tandon, et. al. (Reference 19)

As with the volumes, three junction flow regime maps are used:

• horizontal map for flow in pipes • vertical map for flow in pipes/bundles • high mixing map for flow in pumps

The vertical flow regime map is for junctions whose junction inclination (vertical) angle 𝜙 is such that 60 ≤ |𝜙|≤90 degrees. The horizontal flow regime map is for junctions

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whose junction inclination (vertical) angle 𝜙 is such that 0≤𝜙≤30 degrees. An interpolation region between vertical and horizontal flow regimes is used for junctions whose junction inclination (vertical) angle 𝜙j is such that 30 < 𝜙<60 degrees. This interpolation region is used to smoothly change between vertical and horizontal flow regimes.

Junction quantities used in the map decisions are junction phasic velocities, donored (based on phasic velocities) phasic densities, and donored (based on superficial mixture velocity) surface tension.

∗ The junction void fraction 𝛼, is calculated from either of the volume void fractions of the neighboring volumes, 𝛼, or 𝛼,, using a donor direction based on the mixture superficial velocity (𝑗).

6.2.4 Momentum Closure Relations

NRELAP5 uses two different models for the phasic interfacial friction force computation, the drift flux method and the drag coefficient method. The choice of which model to use depends upon the flow regime.

6.2.4.1 Drift Flux Model

The drift flux approach is used only in the bubbly and slug-flow regimes for vertical flow. The drift flux model specifies the distribution coefficient and the vapor/gas drift velocity. These two quantities must be converted into a constitutive relation for the interfacial frictional force per unit volume.

Such a relation can be found by assuming that the interfacial friction force per unit volume is given by Equation 6-13.

Fi =𝐶𝑖|v𝑅|v𝑅 =𝛼𝛼𝜌 −𝜌𝑔 Equation 6-13

where the interfacial frictional force per unit volume is balanced by the buoyancy force per unit volume where 𝐶 is an unknown coefficient and v is the relative velocity between the phases. Within the context of the drift flux model, the relative velocity between the phases is not the difference between the phasic velocities but is a weighted difference between the phase velocities given by Equation 6-14.

v𝑅 =𝐶1v𝑔 −𝐶0v𝑓 Equation 6-14

where 𝐶 is given by the drift flux correlations and 𝐶 is given by Equation 6-15.

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1−𝛼𝑔𝐶0 𝐶1 = Equation 6-15 1−𝛼𝑔

Substituting these relations into Equation 6-13 gives the interfacial friction force per unit volume in terms of the phasic velocities, given by Equation 6-16.

𝐹𝑖 =𝐶𝑖𝐶1v𝑔 −𝐶0v𝑓𝐶1v𝑔 −𝐶0v𝑓 Equation 6-16

Here the coefficient 𝐶 is yet undetermined. The drift flux model also specifies that the relative velocity (v) can be written as the ratio of the vapor/gas drift velocity and the liquid volume fraction, and is given by Equation 6-17.

v𝑔𝑗 v𝑅 = Equation 6-17 𝛼𝑓

where the vapor/gas drift velocity v is given by the drift flux correlations. Substituting this value of the relative velocity into Equation 6-13 allows the coefficient 𝐶 to be determined from Equation 6-18.

𝛼 𝛼3 𝜌 −𝜌 𝑔 𝑔 𝑓 𝑓 𝑔 Equation 6-18 𝐶i = 2 v𝑔𝑗

6.2.4.2 Drag Coefficient Model

The drag coefficient approach is used in all flow regimes other than vertical bubbly and slug-flow. The model uses correlations for drag coefficients and for the computation of the interfacial area density.

The constitutive relation for the frictional force on a body moving relative to a fluid is given by Equation 6-19.

1 Equation 6-19 F= 𝜌v𝐶 𝐴 2

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where,

𝐹 = drag force

𝜌 = fluid density,

v = velocity of body relative to the fluid,

𝐶 = drag coefficient, and

𝐴 = projected area of the body.

Expressing the frictional force for a group of bodies moving relative to a fluid (e.g., bubbles moving through liquid or droplets moving through vapor/gas) in terms of the frictional force for each body leads to the constitutive relation of Equation 6-20 for the interfacial frictional force per unit volume:

1 F = 𝜌 v −v v −v 𝐶 𝑆 𝑎 =𝐶v −v v −v Equation 6-20 i 8 𝑐 𝑔 𝑓 𝑔 𝑓 𝐷 𝐹 𝑔𝑓 𝑖 𝑔 𝑓 𝑔 𝑓

where,

𝐹 = interfacial friction force per unit volume,

𝐶 = 𝜌 𝐶 𝑆 𝑎

𝜌 = density of continuous phase

𝑎 = interfacial area per unit volume, and

𝑆 = shape factor.

The additional factor of 1/4 comes from the conversion of the projected area of spherical particles (i.e., 𝜋𝑟) into the interfacial area (i.e., 4𝜋𝑟) and the shape factor is included to account for non-spherical particles. The drag coefficient model for the global interfacial friction coefficient has been reduced to the specification of the continuous density, drag coefficient, interfacial area density, and shape factor for the flow regimes. Once these quantities have been computed, the interfacial friction force per unit volume (𝐹) is computed from Equation 6-20 from which the global interfacial friction coefficient can be computed.

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6.2.4.3 Wall Friction

The wall friction is determined based on the volume flow regime map. The wall friction force terms include only wall shear effects. Losses due to abrupt area change are calculated using mechanistic form-loss models. Other losses due to elbows or complicated flow passage geometry are modeled using energy-loss coefficients that must be input by the user.

The semi-implicit scheme, one-dimensional, finite difference equations for the sum momentum equation and the difference momentum equation contain the terms of Equation 6-21 that represent the phasic wall frictional pressure drop.

𝐹𝑊𝐺 ⋅v Δ𝑥 Δ𝑡and 𝐹𝑊𝐹 ⋅v Δ𝑥 Δ𝑡 Equation 6-21

These terms represent the pressure loss due to wall shear from cell center to cell center of the cell volumes adjoining the particular junction that the momentum equation is considering. The wall drag or friction depends not only on the phase of the fluid, but also on the flow regime characteristics.

The wall friction model is based on a two-phase multiplier approach in which the two- phase multiplier is calculated from the heat transfer and fluid flow service (HTFS) modified Baroczy correlation. The individual phasic wall friction components are calculated by apportioning the two-phase friction between the phases using a technique derived from the Lockhart-Martinelli model (Reference 20). The model is based on the assumption that the frictional pressure drop may be calculated using a quasi-steady form of the momentum equation, as used by Chisholm. This wall friction partitioning model is used with the drag coefficient method of the interphase friction model.

The Lockhart-Martinelli model computes the overall two-phase friction pressure drop in terms of the liquid-alone and vapor/gas-alone wall friction pressure drop as shown in Equation 6-22.

𝑑𝑃 𝑑𝑃 𝑑𝑃 =𝜙 =𝜙 Equation 6-22 𝑑𝑥 𝑑𝑥 𝑑𝑥

Here 𝜙 and 𝜙 are the liquid-alone and vapor/gas-alone two-phase Darcy-Weisbach friction multipliers, respectively. The phasic wall friction pressure gradients are expressed by Equation 6-23 for the liquid and vapor/gas alone.

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= and = Equation 6-23

Here the prime indicates the liquid and vapor/gas-alone Darcy-Weisbach friction factors, respectively, calculated at the respective Reynolds numbers given by Equation 6-24.

𝑅𝑒 = and 𝑅𝑒 = Equation 6-24

The liquid and vapor/gas mass flow rates, respectively, are defined by Equation 6-25.

𝑀 =𝛼𝜌v𝐴 and 𝑀 =𝛼𝜌v𝐴 Equation 6-25

The overall two-phase friction pressure gradient is calculated using two-phase friction multiplier correlations. The multipliers are interrelated using Equation 6-22 and Equation 6-23 and the Lockhart-Martinelli ratio defined by Equation 6-26.

𝜒 = = Equation 6-26

The HTFS correlation is used to calculate the two-phase friction multipliers. This correlation was chosen because it is correlated to empirical data over broad ranges of phasic volume fractions, phasic flow rates and phasic flow regimes. The correlation has also been shown to give good agreement with empirical data.

The HTFS correlation for the two-phase friction multiplier is expressed with Equation 6-27.

𝜙 =1+ + and 𝜙 = 𝜒 +𝐶𝜒+1 Equation 6-27

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𝐶 is the correlation coefficient and 𝜒 is the Lockhart-Martinelli ratio given by Equation 6-26. If the HTFS correlation is combined with the wall friction formulations by combining Equation 6-22 and Equation 6-23, Equation 6-25 and Equation 6-26, and Equation 6-27, then the combined two-friction pressure drop is expressed by Equation 6-28.

=𝜙 =𝜙 = 𝜆𝜌𝛼v + Equation 6-28 𝐶 𝜆𝜌𝛼v 𝜆𝜌𝛼v +𝜆𝜌𝛼v

The phasic wall friction coefficients are defined by Equation 6-29 and Equation 6-30.

𝑑𝑃 𝑍 Equation 6-29 𝐹𝑊𝐹𝛼𝜌v𝐴 = 𝜏 𝑝 =𝛼 𝐴 𝑑𝑥 𝛼 +𝛼𝑍

𝑑𝑃 1 𝐹𝑊𝐺(𝛼𝜌v)𝐴=𝜏𝑝 =𝛼 𝐴 Equation 6-30 𝑑𝑥 𝛼 +𝛼𝑍

Here 𝑍 is defined by Equation 6-31.

𝛼 𝜆𝑅𝑒𝜌v 𝛼 𝑍 = Equation 6-31 𝛼 𝜆𝑅𝑒𝜌v 𝛼

Taking the sum of these two equations gives the overall quasi-static, two-phase wall friction pressure gradient as shown by Equation 6-32.

𝑑𝑃 𝐹𝑊𝐹𝛼𝜌v𝐴 + 𝐹𝑊𝐺(𝛼𝜌v)𝐴= 𝐴 Equation 6-32 𝑑𝑥

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The phasic friction factors used in the wall friction model are computed from correlations for laminar and turbulent flows with interpolation in the transition regime. The friction factor model is simply an interpolation scheme linking the laminar, laminar-turbulent transition, and turbulent flow regimes. The laminar friction factor is calculated by Equation 6-33.

64 𝜆 = 𝑓𝑜𝑟 0 ≤ 𝑅𝑒 ≤ 2,200 Equation 6-33 𝑅𝑒𝜙

Here 𝜙 is a user-input shape factor for non-circular flow channels (𝜙 is 1.0 for circular channels).

The friction factor in the transition region between laminar and turbulent flows is computed by reciprocal interpolation with Equation 6-34.

8,250 𝜆 = 3.75 − 𝜆 −𝜆 +𝜆 , 𝑅𝑒 , , , Equation 6-34

𝑓𝑜𝑟 2,200 < 𝑅𝑒 < 3,000

Here 𝜆, is the laminar factor at a Reynolds number of 2,200, 𝜆, is the turbulent friction factor at a Reynolds number of 3,000, and the interpolation factor is defined to lie between zero and one.

The turbulent friction factor is given by the Zigrang-Sylvester approximation (Reference 22) to the Colebrook-White correlation (Reference 23) with Equation 6-35, where 𝜀 is the surface roughness.

1 𝜀 2.51 𝜀 21.25 =−2log + 1.14 − 2 log + . 𝜆 3.7𝐷 𝑅𝑒 𝐷 𝑅𝑒 Equation 6-35

𝑓𝑜𝑟 𝑅𝑒 ≥ 3,000

6.2.5 Heat Transfer

The liquid and vapor/gas energy solutions include the wall heat flux to liquid or vapor/gas. During boiling, the saturation temperature based on the total pressure is the

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reference temperature, and during condensation the saturation temperature based on the partial pressure is the reference temperature. The general expression for the total wall heat flux is defined by Equation 6-36:

𝑞" =ℎ𝑇 −𝑇+ℎ𝑇 −𝑇+ℎ𝑇 −𝑇 +ℎ𝑇 −𝑇+ℎ𝑇 −𝑇 Equation 6-36

where,

ℎ = heat transfer coefficient to vapor/gas, with the vapor/gas temperature as the reference temperature (W/m2 K),

ℎ = heat transfer coefficient to vapor/gas, with the saturation temperature based on the total pressure as the reference temperature (W/m2 K),

ℎ = heat transfer coefficient to vapor/gas, with the saturation temperature based on the vapor partial pressure as the reference temperature (W/m2 K),

ℎ = heat transfer coefficient to liquid, with the liquid temperature as the reference temperature (W/m2 K),

ℎ = heat transfer coefficient to liquid, with the saturation temperature based on the total pressure as the reference temperature (W/m2 K),

𝑇 = wall surface temperature (K),

𝑇 = vapor/gas temperature (K),

𝑇 = liquid temperature (K),

𝑇 = saturation temperature based on the total pressure (K), and

𝑇 = saturation temperature based on the partial pressure of vapor in the bulk (K).

A boiling curve is used in NRELAP5 to govern the selection of the wall heat transfer correlations when the wall surface temperature is above the saturation temperature (superheated relative to the saturation temperature based on total pressure). When a hydraulic volume is voided and the adjacent surface temperature is subcooled, vapor condensation on the surface is predicted. If non-condensable gases are present, the phenomena are more complex because condensation is based on the partial pressure of vapors present in the region. When the wall temperature is less than the saturation temperature based on total pressure, but greater than the saturation temperature based on vapor partial pressure, a convection condition exists. Figure 6-2 illustrates these three regions of the curve.

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Boiling region CHF point

Nucleate Transition Film Heat flux

[T - T ] [Tspp - Tw] w spt Condensing region

Convection region

Figure 6-2. NRELAP5 boiling and condensing curves

The boiling curve uses the Chen boiling correlation (Reference 24) up to the CHF point.

NRELAP5 will issue a message and stop running if CHFR reduces below one for heat conductors that are in the core. Post-CHF heat transfer is allowed to occur on surfaces outside the core. For instance, post-CHF heat transfer can occur on the outside of the RPV where boiling occurs in the pool of liquid that accumulates in the CNV. Post-CHF heat transfer could also occur on the SG tube surfaces, depending on local conditions.

6.3 Heat Structure Models

Heat structures provided in NRELAP5 permit calculation of the heat transfer across solid boundaries of hydrodynamic volumes. Modeling capabilities of heat structures are general and include fuel pins or plates with nuclear or electrical heating, heat transfer across SG tubes, and heat transfer from pipe and vessel walls. Temperatures and heat transfer rates are computed from the one-dimensional form of the transient heat conduction equation.

Heat structures are represented using rectangular, cylindrical, or spherical geometry. Surface multipliers are used to convert the unit surface of the one-dimensional calculation to the actual surface of the heat structure. Temperature-dependent and space-dependent thermal conductivities and volumetric heat capacities are provided in tabular or functional form either from built-in or user-supplied data.

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Finite differences are used to advance the heat conduction solutions. Each mesh interval may contain different mesh spacing, a different material, or both. The spatial dependence of the internal heat source, if any, may vary over each mesh interval. The time-dependence of the heat source can be obtained from reactor kinetics, one of several tables of power versus time, or a control system variable. Boundary conditions include symmetry or insulated conditions; a heat transfer correlation package; and tables of surface temperature versus time, heat flux versus time, heat transfer coefficient versus time, and heat transfer coefficient versus surface temperature.

The heat transfer correlation package can be used for heat structure surfaces connected to hydrodynamic volumes. The heat transfer correlation package contains correlations for convective, nucleate boiling, transition boiling, and film boiling heat transfer from the wall to the fluid, and it contains reverse heat transfer from the fluid to the wall including correlations for condensation (see Section 6.2.5 and Section 6.8). The heat conduction model also includes a gap conduction model and a radiation enclosure model.

The integral form of the heat conduction equation is defined by Equation 6-37.

𝜕𝑇 𝜌𝐶(𝑇,̅ 𝑥) (𝑥̅,𝑡)𝑑𝑉 = 𝑘(𝑡,̅ 𝑥)∇ 𝑇(𝑡,̅ 𝑥) ⋅𝑑𝑠̅ +𝑆(𝑥̅,𝑡)𝑑𝑉 𝜕𝑡 Equation 6-37

where, 𝑘(𝑡,̅ 𝑥) = thermal conductivity, 𝑠 = surface, 𝑆 = internal volumetric heat source, 𝑡 = time, 𝑇 = temperature, 𝑉 = volume, 𝑥 = space coordinates, and

𝜌𝐶 = volumetric heat capacity.

The boundary conditions applied to the exterior surface have the form of Equation 6-38.

𝜕𝑇(𝑡) 𝐴(𝑇)𝑇(𝑡) +𝐵(𝑇) =𝐷(𝑇,) 𝑡 Equation 6-38 𝜕𝑛

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The 𝑛 denotes the unit normal vector away from the boundary surface. Thus, if the desired boundary condition is that the heat transferred out of the surface equals a heat transfer coefficient (ℎ) times the difference between the surface temperature (𝑇) and the sink temperature (𝑇) as shown by Equation 6-39.

𝜕𝑇(𝑡) −𝑘 =ℎ(𝑇−𝑇 ) Equation 6-39 𝜕𝑛

then the correspondence between the above expression and Equation 6-38 yields 𝐴= ℎ, 𝐵 = 𝑘, 𝑎𝑛𝑑 𝐷=ℎ𝑇 .

One-dimensional heat conduction in rectangular, cylindrical, and spherical geometry can be used to represent the heat structures in any of the components in NRELAP5. The equations governing one-dimensional heat conduction are defined by Equation 6-40, Equation 6-41, and Equation 6-42.

𝜌𝐶 = 𝑘 +𝑆 for rectangular geometry Equation 6-40

𝜌𝐶 = 𝑟𝑘 +𝑆 for cylindrical geometry Equation 6-41

𝜌𝐶 = 𝑟𝑘 +𝑆 for spherical geometry Equation 6-42

Heat may flow across the external heat structure boundaries to either the environment or to the working fluid. For heat structure surfaces connected to hydrodynamic volumes containing the working fluid, a heat transfer package is provided containing correlations for heat transfer from wall-to-liquid and reverse heat transfer from liquid-to-wall. Any number of heat structures may be connected to each hydrodynamic volume, or heat transfer coefficient versus surface temperature can be used to simulate the boundary conditions.

The heat conduction equation can be solved by various numerical techniques. NRELAP5 uses the Crank-Nicolson method (Reference 26) for solving this equation.

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6.4 Point Reactor Kinetics Model

NRELAP5 allows the user to model the power generated in the reactor core as specified from a table or as determined by point-reactor kinetics with reactivity feedback. This power is modeled as an internal heat source in user-defined heat structures and can be partitioned by inputting weighting factors to distribute the energy to the various portions of the core as the user desires.

The power is computed using the space-independent or point kinetics approximation, which assumes that power can be separated into space and time functions. The point reactor kinetics model computes both the immediate (prompt and delayed neutrons) fission power and the power from decay of fission products. The immediate (prompt and delayed neutrons) power is that released at the time of fission and includes power from kinetic energy of the fission products and neutron moderation. Decay power is generated as the fission products undergo radioactive decay. The LOCA methodology uses the ANS 1973 decay heat standard (see Section 6.10).

The point kinetics equations are (see Glasstone and Sesonske, Reference 27) defined by

𝑑𝑛(𝑡) 𝜌(𝑡) −𝛽 = 𝑛(𝑡) +𝜆𝐶 (𝑡) +𝑆 Equation 6-43 𝑑𝑡 Λ

𝑑𝐶 (𝑡) 𝛽𝑓 = 𝑛(𝑡) −𝜆𝐶 (𝑡) 𝑖=1,2,…,𝑁 Equation 6-44 𝑑𝑡 Λ

𝜑(𝑡) =𝑛(𝑡)𝑣 Equation 6-45

Ψ(𝑡) =𝑉Σ𝜑(𝑡) Equation 6-46

𝑃(𝑡) =𝑄Ψ(𝑡) Equation 6-47

where, t = time (s),

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n = neutron density (neutrons/m3), ϕ = neutron flux (neutrons/m2·s), v = neutron velocity (m/s),

3 Ci = delayed neutron precursor concentration in group i (nuclei/m ),

β = effective delayed neutron fraction = ∑ β, Λ = prompt neutron generation time (s) ρ = reactivity (only the time-dependence has been indicated; however, the reactivity is dependent on other variables),

fi = fraction of delayed neutrons of group i = βi/β,

βi = effective delayed neutron precursor yield of group i,

λι = decay constant of group i (1/s), S = source rate density (neutrons/m3·s), ψ = fission rate (fissions/s), Σ f = macroscopic fission cross-section (1/m),

Pf = immediate (prompt and delayed neutron) fission power (MeV/s),

Qf = immediate (prompt and delayed neutron) fission energy per fission (MeV/fission), V = volume (m3), and

Nd = number of delayed neutron precursor groups.

After some modifications and variable substitutions, these equations are solved by the modified Runge-Kutta method.

Reactivity feedback can be input into NRELAP5 in one of two models: a separable model and a tabular model. The separable model is so defined that it assumes that each effect is independent of the other effects. This model also assumes nonlinear feedback effects from moderator density and fuel temperature changes and linear feedback from moderator temperature changes.

6.5 Trips and Control System Models

The control system provides the capability to evaluate simultaneous algebraic and ordinary differential equations. The capability is primarily intended to simulate control systems typically used in hydrodynamic systems, but it can also model other phenomena described by algebraic and ordinary differential equations. Another use is to define auxiliary output quantities, such as differential pressures, so they can be printed in major and minor edits and be plotted.

The control system consists of several types of control components. Each component defines a control variable as a specific function of time-advanced quantities. The time-

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advanced quantities include hydrodynamic volume, junction, pump, valve, heat structure, reactor kinetics, trip quantities, and the control variables themselves (including the control variable being defined). This permits control variables to be developed from components that perform simple, basic operations.

The trip system consists of the evaluation of logical statements. Each trip statement is a simple logical statement that has a true or false result and an associated variable. Two types of trip statements are provided (variable and logical trips).

6.6 Special Solution Techniques

Certain models in NRELAP5 have been developed to simulate special processes. Special process models are used in NRELAP5 to model those processes, which are sufficiently complex that they must be modeled by empirical models. The following sections summarize those models.

6.6.1 Choked Flow

6.6.1.1 Moody Critical Flow Model

Because the Moody model (Reference 28) is required by 10 CFR 50 Appendix K when the break flow is calculated to be two-phase, a critical flow model that complies with the 10 CFR 50 Appendix K requirements was incorporated in NRELAP5. Two options are available in NRELAP5 for use of the Moody model. {{

}}2(a),(c)

Moody developed his critical flow model from theory to predict the maximum flow rate of a single component, two-phase mixture. The model assumes that the liquid phase is incompressible and that the flow is isentropic so that the stagnation enthalpy is constant throughout the system. The flow is maximized with respect to local slip ratio and static pressure for known stagnation conditions. The specific volume (𝜐) and specific enthalpy (h) of water can be calculated from two state variables, entropy (𝑠) and pressure (P), i.e.,

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ℎ=ℎ(𝑠,𝑃) and 𝜐=𝜐(𝑠,𝑃), where the subscript 0 denotes break entrance conditions. Because entropy is constant, ℎ and 𝜐 are functions of P, the stagnation pressure. From the continuity and energy equations for homogeneous flow entering and leaving an ideal nozzle the mass flux, G, satisfies:

/ 𝐺=2(ℎ −ℎ)/𝜐 Equation 6-48

The maximum flow rate occurs at the throat, where

𝑑𝐺 Equation 6-49 | =0 𝑑𝑃

1/3 Moody showed that the maximum flow occurs when the slip ratio K = (vg/vf) . With this value of the slip ratio, Moody derived a complex equation for the critical flow rate that was used to create Moody lookup tables for the flow rate as a function of stagnation pressure and stagnation enthalpy. The range of the tables that are used in NRELAP5 covers local static pressure from {{ }}2(a),(c) with local quality from 0.0 to 1.0 and local stagnation pressures and enthalpies covering the range of saturation states. {{

}}2(a),(c)

6.6.1.2 Henry-Fauske Critical Flow Model

The principle assumption used in the Henry-Fauske model is that, for most applications, the amount of thermal non-equilibrium at the throat is more important in determining the critical flow rate than the amount of mechanical non-equilibrium. Thus, it is assumed that the phase velocities are equal. Henry and Fauske then argued that for normal nozzle configurations, there is little time for mass transfer to take place, and it is reasonable to assume that the amount of mass transferred during the expansion is negligible and also that the amount of heat transferred between the phases during the expansion is insignificant, so that the liquid temperature is essentially constant. Interfacial viscous terms were neglected. Based on these assumptions Henry and Fauske derived an equation for the mass flux at the throat. The mass flow rate exhibits a maximum with respect to the throat pressure at critical flow, which yields a complex relationship for the critical mass flux that includes dependency on the throat pressure.

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The Henry-Fauske model requires only knowledge of the upstream stagnation conditions and, unlike earlier critical flow models, it accounts for the non-equilibrium nature of the flow. Henry and Fauske noted that the critical flow rates are in reasonable agreement with the homogeneous equilibrium model for stagnation qualities greater than 0.10, and that for qualities less than this value, the homogeneous equilibrium model underestimates the data. Therefore, they required that the model input use only stagnation conditions, and yet at the same time account for the non-equilibrium nature of the flow. To address this issue Henry and Fauske correlated the effect of thermal non- equilibrium on the mass transfer rate at the throat as:

𝑑𝑋 𝑑𝑋 Equation 6-50 | =𝑁 | 𝑑𝑃 𝑑𝑃

where 𝑁 is a thermal non-equilibrium factor defined in terms of the equilibrium quality at the throat (𝑋,).

𝑋, Equation 6-51 𝑁= 0.14

The final remaining unknown is the value of the pressure at the throat. To determine the throat pressure, the two-phase momentum equation was integrated between the stagnation and the throat locations to give an equation for the critical pressure ratio, i.e., the ratio of the throat pressure to the upstream stagnation pressure when the flow is choked. The use of this equation for the throat pressure in the equation for the critical mass flux results in a transcendental equation for the critical mass flux. The solution of the transcendental equation implicitly involves the critical flow rate and hence its solution yields predictions of the critical pressure ratio and the critical flow rate as functions of the upstream stagnation pressure and quality. The critical pressure ratio determines the transition to non-choked flow. If the mass flux predicted by the critical flow model is less than that resulting from the normal solution of the momentum equations, then the junction is choked. Assessment of the Henry-Fauske model shows excellent agreement against the Marviken 22 and 24 tests (Section 7.2.11).

6.6.1.3 Choked Flow for Orifices, Nozzles and Valves

To provide the user with the ability to better characterize the orifice, nozzle or valve behavior, the form of the Henry-Fauske model was retained in RELAP5-3D© and carried over into NRELAP5. The constant in the thermal non-equilibrium factor is included as an adjustable parameter.

𝑋, Equation 6-52 𝑁=𝑀𝐼𝑁(1, ) 𝐶

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where the thermal non-equilibrium constant, 𝐶, is user input with a default value of 0.14. A user input discharge coefficient, default value of 1.0, can also be applied to the critical mass flux. The ability to input these parameters allows the user to adjust the critical flow model to account for the different amount of thermal non-equilibrium at the throat.

While the model development was based on a converging nozzle, the authors of the model extended the results to orifices and short tubes by comparison to experimental data for these geometries. The Henry-Fauske model can be applied to cases where the upstream condition is subcooled liquid or single-phase vapor. While the Henry-Fauske model can handle non-condensable gas, the total amount of non-condensable gas in the NPM is negligible so this capability is not addressed in the following discussion.

During the development of the RELAP5 codes, modifications were made to the original model to ensure continuity at phase transitions to better characterize nozzles and orifices. Specifically,

• the phase transition modifications provide a smooth transition of the critical flow at the subcooled liquid to two-phase mixture interface. • two adjustable coefficients, a discharge coefficient and a thermal non-equilibrium constant are provided in order to better characterize nozzles and orifices. The discharge coefficient is a multiplier on the flow area. The non–equilibrium constant is an assumed throat equilibrium quality that was assigned an average value of 0.14 by Henry and Fauske, but can be specified by the code user.

With these modifications, the Henry-Fauske model is applicable to two-phase and single-phase superheated and subcooled critical flow. The two adjustable coefficients allow the code user to more closely match test data from the valve vendor and calibration data from orifices and nozzles used in experimental facilities.

6.6.2 Abrupt Area Change

The general reactor system contains piping networks with many sudden area changes and orifices. To apply the NRELAP5 hydrodynamic model to such systems, analytical models for these components are included in the code. The basic hydrodynamic model is formulated for slowly varying (continuous) flow area variations; therefore, special models are not required for this case.

The abrupt area change model, is based on the Borda-Carnot formulation (Reference 30) for a sudden (i.e., sharp, blunt) enlargement and standard pipe flow relations, including the vena-contracta effect for a sudden (i.e., sharp, blunt) contraction or sharp- edge orifice or both. This is referred to as the full abrupt area change model. It does not include the case where an enlargement, contraction, or orifice is rounded or beveled.

Quasi-steady continuity and momentum balances are employed at points of an abrupt area change. The numerical implementation of these balances is such that hydrodynamic losses are independent of upstream and downstream nodalization. In effect, the quasi-steady balances are employed as jump conditions that couple fluid

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components having abrupt changes in cross-sectional area. This coupling process is achieved without change to the basic numerical time-advancement schemes.

The basic assumption used for the transient calculation of two-phase flow in flow passages with points of abrupt area change is that the transient flow process can be approximated as a quasi-steady flow process that is instantaneously satisfied by the upstream and downstream conditions (that is, transient inertia, mass, and energy storage are neglected at abrupt area changes). However, the upstream and downstream flows are treated as fully transient flows.

The volume of fluid and associated mass, energy, and inertia at points of abrupt area change is generally small compared with the volume of upstream and downstream fluid components. The transient mass, energy, and inertia effects are approximated by lumping them into upstream and downstream flow volumes. Finally, the quasi-steady approach is consistent with modeling of other important phenomena in transient codes (that is, heat transfer, pumps, and valves).

Activation of the full abrupt area change model in NRELAP5 results in the code internally calculating the form and interfacial losses across a junction. Utilization of the partial area change model allows the user to specify the form loss while allowing the code to internally calculate the interfacial loss. Activation of the smooth area change model allows the user to specify the form loss with no internal calculation of the interfacial losses.

More detailed discussion concerning this model can be found in the NRELAP5 theory manual (Reference 9).

6.6.3 Counter Current Flow Limitation

A general CCFL model is implemented in a form proposed by Bankoff (Reference 31), which has the structure

/ / 𝐻 +𝑚𝐻 =𝑐 Equation 6-53

where,

𝐻 = dimensionless vapor/gas flux,

𝐻 = dimensionless liquid flux,

/ 𝑐 = vapor/gas intercept (value of 𝐻 when 𝐻 =0, i.e., complete flooding), and

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m = “slope”, that is the vapor or gas intercept divided by the liquid intercept (the value / of 𝐻 when 𝐻 =0).

The dimensionless fluxes have the form as defined by Equation 6-54 and Equation 6-55.

/ 𝜌 𝐻 =𝑗 Equation 6-54 𝑔𝑤𝜌 −𝜌

/ 𝜌 𝐻 =𝑗 Equation 6-55 𝑔𝑤𝜌 −𝜌

In these equations 𝑗 is the vapor/gas superficial velocity 𝛼𝑣, 𝑗 is the liquid superficial velocity 𝛼𝑣, 𝜌is the vapor/gas density, 𝜌 is the liquid density, 𝛼 is the vapor/gas volume fraction, 𝛼 is the liquid volume fraction, g is the gravitational acceleration. In Equation 6-54 and Equation 6-55, 𝑤 is the length scale and is given by Equation 6-56.

𝑤=𝐷 𝐿 Equation 6-56

Where 𝛽 is a user-input constant, 𝐷 is the junction hydraulic diameter and 𝐿 is the Laplace capillary length constant given by Equation 6-57.

𝜎 / 𝐿= Equation 6-57 𝑔𝜌 −𝜌

Bankoff recommends a formula for computing the value of 𝛽:

𝐴 2𝜋 𝛽=tanh ∙ ∙𝐷 Equation 6-58 𝐴 𝑡

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where 𝐴 is the total area of the holes through the plate, 𝐴 is the total area of the plate, including the holes, and 𝑡 is the thickness of the plate.

The Bankoff correlation specifies that the vapor/gas intercept (c) is of the form:

𝑐=(1.07 + 0.00433 ∙ 𝐿) Equation 6-59

when the dimensionless Bond number, 𝐿, is less than 200 and 𝑐=2 for all 𝐿 greater than or equal to 200.

The bond number is:

/ 𝑔𝜌 −𝜌 𝑛𝜋𝐷 𝐿 =𝑛𝜋𝐷 = Equation 6-60 𝜎 𝐿

where 𝑛 is the number of holes in the plate at the CCFL junction, 𝐷 is the hydraulic diameter, and 𝐿 is the Laplace capillary length constant, previously defined. More detailed discussion concerning this model can be found in the NRELAP5 theory manual (Reference 9).

Assessment of the CCFL model demonstrates excellent agreement against Bankoff perforated plate test data (Section 7.2.10). A sensitivity study of the effects of the CCFL model is presented in Section 9.6.3 for its application at the NPM pressurizer baffle plate.

6.7 Helical Coil Steam Generator Component

A new hydrodynamic component and heat transfer package have been added to NRELAP5 to model flow and heat transfer inside a helical coil SG. These are developed based on helical coil geometry-specific heat transfer and wall friction correlations. The need for improved models is based on inadequate agreement with pressure drop and heat transfer performance with the baseline RELAP5-3D© code results against prototypic helical coil SG testing performed at SIET. Improvements and adequacy of the implemented models in NRELAP5 are demonstrated through prototypic assessments of the NuScale helical coil SG using SIET test data (see Sections 7.4.1 and 7.4.2). These tests assessed heat transfer and pressure drop on both the secondary side (within tubes) and primary side (external to tubes) of the helical coil SG.

A wide range of pressure drop and heat transfer correlations were investigated for analyzing the inside of the helical coils. A down selection was performed of these

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investigated models for implementation into the NRELAP5 code based on the applicability of the models to the NPM helical coil SG.

6.7.1 Helical Coil Tube Friction

6.7.1.1 Helical Coil Single-Phase Tube Wall Friction

The {{ }}2(a),(c) provided the best global coverage and as such have been implemented into NRELAP5. {{

}}2(a),(c)

6.7.1.2 Helical Coil Two-Phase Tube Wall Friction

The two-phase inner wall friction for a helical coil is computed in a similar fashion to the Lockhart-Martinelli model implemented in the RELAP5 code series. A modification is made to the two-phase friction multiplier for the helical coil component as presented in its final form by Equation 6-62.

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{{

}}2(a),(c)

6.7.2 Helical Coil Tube Heat Transfer

A new heat transfer package has been added to NRELAP5 and differs from that of the standard RELAP5 pipe geometry in {{ }}2(a),(c) A new geometry type represents the {{

}}2(a),(c)

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6.7.2.1 Helical Coil Single-Phase Heat Transfer

The laminar heat transfer correlation {{

}}2(a),(c)

6.7.2.2 Helical Coil Two-Phase Subcooled and Saturated Flow Boiling Heat Transfer

The saturated flow boiling heat transfer correlation is used for {{

}}2(a),(c)

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{{

}}2(a),(c)

6.8 Wall Heat Transfer and Condensation

Due to the significance of CNV wall heat transfer in reactor core cooling and decay heat removal during a postulated NPM LOCA (see Section 8.2.8), a detailed discussion is presented in this section on NRELAP5 wall heat transfer and condensation models. {{

}}2(a),(c)

As described below, the {{

}}2(a),(c)

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{{

}}2(a),(c)

Section 6.8.1 below provides further discussion on NRELAP5 evaluation of wall heat transfer with film condensation. The discussion includes the definition of the liquid (film) Reynolds number, partitioning of the total wall heat flux between liquid and vapor phases, and handling the effect of non-condensable gases that may be present in the hydrodynamic volume. Section 6.8.2 summarizes the extended Shah correlation used in NRELAP5 for wall condensation.

6.8.1 NRELAP5 Solution Approach for Wall Condensation Heat Transfer

NRELAP5 solves {{

}}2(a),(c)

{{

}}2(a),(c)

Figure 6-3. {{ }}2(a),(c)

{{ }}2(a),(c)

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{{

}}2(a),(c)

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{{

}}2(a),(c)

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{{

}}2(a),(c)

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6.8.2 Wall Condensation Correlation

{{

}}2(a),(c)

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{{

}}2(a),(c)

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Table 6-1. Extended Shah dimensionless vapor velocity transition criteria {{

}}2(a),(c) Table 6-2. Extended Shah condensation heat transfer coefficients dependent on regime {{

}}2(a),(c)

6.9 Interfacial Drag in Large Diameter Pipes

RELAP5-3D© contains the Kataoka-Ishii (Reference 40) formulation of the drift-flux model for use in the bubbly flow case in intermediate (0.018 < 𝐷 ≤ 0.08 𝑚) and large pipes (𝐷 > 0.08 𝑚). This same dimensional formulation is maintained within NRELAP5.

RELAP5-3D© originally implemented the modified Rouhani distribution coefficient (Reference 42) as shown by Equation 6-87.

𝜌𝑔𝐷 𝐶 =1+0.2 Equation 6-93 |𝐺| + 0.001

{{

}}2(a),(c)

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6.10 Fission Decay Heat and Actinide Models

The ANS 1973 fission decay heat standard (Reference 46) is presented in terms of the Shure curve (Reference 47) and tabular data. The NRELAP5 implementation of the ANS 1973 standard applies the Shure curve, which is a fit to differential equations for one isotope and 11 groups. Assuming infinite operating time, the fission product decay power is calculated with Equation 6-89. Table 6-3 provides the 11-group constants derived from the Shure curve as implemented into NRELAP5. Figure 6-3 provides the comparison of the ANS 1973 standard to the as implemented curve.

P=P𝛾𝐴exp (−𝑎𝑡) Equation 6-95

where,

𝑃 = fission decay power,

𝑃 = infinite operating time fission power prior to shutdown,

𝛾 = fission product yield factor,

𝐴,𝑎 = fit coefficients, and

𝑡 = time after shutdown.

Table 6-3. ANS 1973 11-group fission decay constants

{{

}}2(a),(c)

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Figure 6-4. NRELAP5 ANS 1973 implemented fission decay heat curve

The actinide model describes the production of 239U, 239Np, and 239Pu from neutron capture by 238U based on the decay equations of Equation 6-90.

dγ() =𝐹 Ψ(𝑡) −𝜆 𝛾 dt

dγ (𝑡) Equation 6-96 =𝜆 𝛾 −𝜆 𝛾 dt

P(𝑡) =𝜂𝜆𝛾(𝑡) +𝜂𝜆𝛾(𝑡)

239 The quantity FU is user-specified and is the number of atoms of U produced by neutron capture in 238U per fission from all isotopes. The 𝜆 and 𝜂 values can be user- specified, or default values equal to those stated in the 1979 ANS standard (Table 6-4), the 1994 Standard, or the 2005 Standard can be used.

The first equation describes the rate of change of atoms of 239U. The first term on the right represents the production of 239U; the last term is the loss of 239U due to beta decay.

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The second equation describes the rate of change of 239Np. The production of 239Np is from the beta decay of 239U, and 239Pu is formed from the decay of 239Np. Ψ(𝑡) is the solution from the NRELAP5 fission source. The implemented model yields the result quoted in the 1979 Standard (Reference 48), the 1994 Standard (Reference 49), and the 2005 Standard (Reference 50) as demonstrated by Figure 6-4.

Table 6-4. ANS-79 actinide model constants.

Isotope 𝜆(s-1) 𝜂(MeV) 239U 1.772 0.00299 239Np 0.5774 0.00825

Figure 6-5. NRELAP5 ANS-79 implemented actinide heat curve

6.11 Critical Heat Flux Models

The CHF is calculated using a combination of the {{

}}2(a),(c)

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{{

}}2(a),(c)

6.11.1 {{ }}2(a),(c)

{{

}}2(a),(c)

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{{

}} 2(a),(c) Figure 6-6. {{ }}2(a),(c)

{{

}}2(a),(c)

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{{

}}2(a),(c)

6.11.2 Implementation of Critical Heat Flux correlations

{{ }}2(a),(c) are implemented in NRELAP5 as follows:

• {{

}}2(a),(c)

6.11.3 {{ }}2(a),(c)

{{

}}2(a),(c)

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{{

}}2(a),(c)

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{{

}}2(a),(c)

Table 6-5. Coefficient of revised pressure correction term in Equation 6-108 {{

}}2(a),(c)

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{{

}}2(a),(c)

Figure 6-7. {{ }}2(a),(c)

{{

}}2(a),(c)

Table 6-6. {{ }}2(a),(c) critical heat flux correlation application range

{{

}}2(a),(c)

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{{

}}2(a),(c)

6.11.4 {{ }}2(a),(c)

{{

}}2(a),(c)

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{{

}}2(a),(c)

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7.0 NRELAP5 Assessments

The following section provides a summary of the SET and IET assessments that have been completed for NRELAP5. The results of these assessments are considered in Section 8.0 to justify the adequacy on NRELAP5 for modeling of high-ranked phenomena in the NuScale LOCA PIRT.

To assess the adequacy of NRELAP5, code simulations are compared to measured experimental data. Acceptance criterion from Table 1-2 are applied in rating NRELAP5 performance in terms of “excellent”, “reasonable” or “minimal” agreement. These ratings take into consideration the ability to predict overall data trends as well as the magnitude of the data itself.

7.1 Assessment Methodology

Various experimental tests, inclusive of SETs, IETs, and analytic problems have been used to assess the performance of NRELAP5 using the process identified in Element 2 of RG 1.203. The database employed to assess the adequacy of the NRELAP5 code was chosen to be consistent with the requirements to adequately model the high-ranked phenomena derived in the NuScale LOCA PIRT.

The high-ranked phenomena selected in Section 4 are mapped onto an assessment matrix of experiments, and are listed in Table 7-1. The analytic problems (fundamental tests) used to assess NRELAP5 are not shown in Table 7-1.

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Table 7-1. NRELAP5 loss-of-coolant accident assessment matrix {{

}}2(a),(c)

Summarized within this section for each assessment are the following:

• a brief description and purpose of the experimental facility, • a summary of the phenomenon addressed, • the experiment procedure, • important NRELAP5 modeling techniques, and • performance of NRELAP5 against the data.

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Assessment cases are divided into two categories: • Legacy Assessments – these are assessments performed against data collected from historical test programs not encompassed within the NuScale test programs • NuScale Test Assessments – these are assessments performed against data collected as part of the NuScale testing program

The following sections document the various assessments completed with NRELAP5.

7.2 Legacy Test Data

This section describes those test programs which have produced data that were not performed under the NuScale QAPD (Reference 4). With the exception of Marviken JIT- 11 data, these tests are qualified for use by applying non-mandatory guidance provided by NQA-1 2008/1a-2009 Addenda (Reference 10). Use of Marviken JIT-11 data is based on published literature data.

7.2.1 Ferrell-McGee

The Ferrell-McGee tests were performed in vertical pipes over a wide range of single- phase and two-phase flow conditions with uniform, contraction, and expansion flow areas. The data assessed includes single- and two-phase pressure drop and void fraction under different pressures, flow rates, and inlet quality.

7.2.1.1 Facility Description

The report for the Ferrell-McGee experiments (Reference 58) describes the test facility. Figure 7-1 shows the schematic of the test section. The test apparatus consists of a heated section that controls the degree of sub-cooling of the liquid entering into an adiabatic test section, where pressure drops and void fractions were measured. The lower test section is 40.5 in. (1.0287 m) in height and the upper test section is 49.5 in. (1.2573 m) for a total of 90.0 in. (2.286 m). The two test sections were connected by mating flanges. The tests were organized into seven test groups. Each test group had a different combination of pipe diameters for the lower and upper sections. The tests of Group 1 and Group 4 used pipes of uniform diameter of 0.46 in. (0.0117 m) and 0.34 in. (0.00864 m) and are designed to assess two-phase frictional pressure drop. Tests of Groups 2, 5 and 6 are tests with abrupt area expansion with area ratios of 0.608, 0.332 and 0.546, respectively. The tests of Groups 3 and 7 are tests with abrupt contraction with area ratios of 0.546 and 0.608, respectively. In this section the area ratios are defined as small area or large area. Tests with abrupt area expansion and contractions are designed to assess frictional and form losses. Multiple sets of tests were run with different combinations of pressure, flow, and inlet quality.

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Figure 7-1. Schematic of the Ferrell-McGee test section

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7.2.1.2 Phenomena Addressed

The phenomenon addressed with the Ferrell-McGee assessment cases is the ability of NRELAP5 to predict {{ }}2(a),(c)

7.2.1.3 Experimental Procedure

The part of the stainless steel flow loop of particular interest is the vertical adiabatic test section in which an upward flowing steam-water mixture entered at a controlled pressure, mass flow rate, and quality. Pressure drops and steam volume fractions (α) were measured along the channel at the locations shown in Figure 7-1.

The two-phase mixture exited from the test section into a 0.460-in. inside diameter glass section through which the mixture could be photographed. The vapor-liquid mixture was partially separated in a surge tank and sub-cooled in a bank of six parallel concentric- tube heat exchangers.

The sub-cooled liquid passed through a pump, a manual flow control valve, a volumetric flowmeter, and a preheater which controlled the quality of two-phase flow entering the heated section. In the heated section, a 0.462-in. inside diameter by 0.083-in. wall tube heated by an alternating current flowing in the tube wall, the water was brought to the desired quality before injection into the adiabatic test section.

System pressure was maintained by a hydraulic accumulator. Loop fluid, cooling water, heated channel wall, and manometer line temperatures were also recorded.

The summary of ranges of recorded data for the 201 runs is provided in Table 7-2. The initial boundary conditions covered a range of three different mass flow rates of 460, 920 and 1,150 lbm/hr (209, 417, and 522 kg/hr), a range of three different pressures of 60, 120 and 240 psi (0.414, 0.827, and 1.65 MPa) and void fractions from 0.0 to 1.0.

Table 7-2. Summary of Ferrell-McGee experimental test data range

Range Parameter Min Max Units Pressure 60 (0.414) 240 (1.65) psia (MPa) Inlet flow rate 460 (209) 1,150 (522) lbm/hr (kg/hr) Inlet void fraction range1 -0.110 1.038 n/a Expansion area ratio 0.332 0.608 n/a Contraction area ratio 0.546 0.608 n/a 1Negative void fractions refer to sub-cooling estimates which are calculated.

Although the measured data includes void distributions at nine different locations, shown in Figure 7-1, only the void at measurement locations 1, 2 and 3 were considered. These locations have the void measurements near the center of NRELAP5 nodes. The rest of the void fraction measurement taps were not taken into account because placing all void

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measurement locations near the center of nodes would result in small nodes with a length-to-diameter (L/D) ratio less than 1.0.

The total pressure drop measurement uncertainty was estimated to be ±0.45 psi (0.0031 MPa). The average void fraction measurement uncertainty was estimated to be ±3 percent.

It is noted that test cases 1A6, 1A7, 1A11, 4A5, 4A9, 5A5, 5A10 and 6A9 with void fractions at Void Tap 1 larger than 0.97 were excluded from comparative results. The NRELAP5 total pressure drop predictions for these test cases showed a high deviation from the measurement, which subsequent analysis revealed to result from uncertainty of the inlet void fraction measurement.

7.2.1.4 Special Analysis Techniques

For test groups 2, 3, 5, 6 and 7, at the position of expansion and contraction {{

}}2(a),(c)

7.2.1.5 Assessment Results

Figure 7-2 shows the predicted versus measured pressure drop for uniform, expansion, and contraction tests. NRELAP5 predicted the experimental data with reasonable-to- excellent agreement. These results validate the ability of NRELAP5 to predict {{

}}2(a),(c)

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Figure 7-2. Predicted versus measured pressure drop for selected contraction tests

7.2.2 GE Level Swell (1 ft)

During various phases of a blowdown event in the NPM, such as a LOCA, the fluid within the RCS will experience flashing, vapor generation, level swell, and conditions representative of rapid depressurization. Reference 59 produced a suitable experimental database extending across a large range of pressures and fluid conditions that is used to assess the ability of NRELAP5 to predict {{ }}2(a),(c) The assessment of NRELAP5 against the 1 ft. diameter GE level swell test is summarized in this section, while the assessment of NRELAP5 against the 4 ft. GE level swell test is provided in Section 7.2.3.

7.2.2.1 Facility Description

The experimental facility is fully described in Reference 58 and summarized in this section. The experiment facility shown in Figure 7-3 consists of a pressure vessel made of carbon steel with a volume of 10 ft3 (0.283 m3), a diameter of approximately 12 in. (0.305 m) and a 14-ft (4.2672 m) length. The small vessel experiments (1 ft.) include a blowdown line with orifice plates that are interchanged to control the blowdown flow rate and depressurization rate. The effluent from the vessel blowdown is discharged into a suppression tank.

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blowdown 14 ft pressure vessel orifice

12 ft ΔP

10 ft ΔP

8 ft ΔP

ΔP T saturated liquid 6 ft T ΔP

blowdown valve 4 ft T ΔP rupture disc asssembly

2 ft T ΔP

T 0 ft heater connections

1 ft

suppression tank

Figure 7-3. Schematic of the GE 1 ft. blowdown vessel

Three basic types of measurements were obtained during each experiment: static pressures, differential pressures, and temperatures.

Figure 7-3 shows the location of the instrumentation. There are six measurement sections between the adjacent differential pressure taps, numbered sequentially. The two-phase density (or mixture density) in each measurement section during blowdown experiments is derived from the axial differential pressure measurements. The fluid mass inventory is obtained from the density and known volume of the measurement section.

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The average void fraction in each measurement section is determined from the measured mixture density and thermodynamic properties of the liquid and vapor phases at the system pressure as shown in Equation 7-1 (Reference 59).

𝛼 =𝜌̅ −𝜌/𝜌 −𝜌 Equation 7-1

Where 𝛼 is the average void fraction in the i-th measurement section, 𝜌̅ is the average mixture density in the i-th measurement section, and 𝜌,𝜌 are the liquid and vapor densities as a function of the measurement sectionpressure.

7.2.2.2 Phenomena Addressed

The phenomena addressed with the GE level swell (1 ft. diameter) test include {{

}}2(a),(c)

Specifically, the GE level swell test assesses the ability of NRELAP5 to predict key in- vessel thermal-hydraulic phenomena associated with a rapid depressurization event.

7.2.2.3 Experimental Procedure

The experiment consisted of filling the vessel with demineralized water and boiling the inventory at atmospheric pressure to remove any dissolved gas. The top vent was then closed and the fluid was heated to the specified initial conditions, which was typically saturated conditions at the desired pressure. The initial water level was dependent upon the experiment of interest as listed in Table 7-3. Once conditions were reached a blowdown was initiated from a discharge valve located at the top of the vessel and measurements were recorded.

Table 7-3 summarizes the experiment conditions for the GE 1 ft. level swell test selected for the assessment. The parameters listed are used as boundary conditions in the NRELAP5 inputs or for comparisons to the NRELAP5 predictions. The blowdown experiment is initiated at a pressure of 1,011 psia (6.97 MPa) with saturated fluid conditions. The blowdown and fluid response is measured for approximately 300 seconds. Void fractions ranging from 0.0 to 1.0 are present during the test.

Table 7-3. Summary of GE 1 ft. vessel level swell experiments

Test Orifice Size Restriction Plate Initial Pressure Initial Liquid Level Number inches (mm) Configuration psia (MPa) ft (m) 1004-3 3/8 (9.525) No plate 1,011 (6.971) 10.4 (3.17)

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7.2.2.4 Special Analysis Techniques

Based on sensitivity studies it was determined that applying the {{ }}2(a),(c) improved the depressurization comparisons from reasonable-to-excellent agreement. The discussion here, however, provides a summary with default discharge coefficient input of {{ }}2(a),(c)

7.2.2.5 Assessment Results

Figure 7-4 through Figure 7-8 present the vessel pressure and axial void fraction comparisons between NRELAP5 using the {{ }}2(a),(c) choking model and the measured data for experiment 1004-3. The initial 100 seconds of the simulations are to confirm a steady state condition. The blowdown event is initiated at 100 seconds and is therefore the initial time for all figures.

Figure 7-4 presents the calculated vessel pressure versus the measured data. The comparisons are in reasonable-to-excellent agreement. The predicted depressurization rate is slightly higher compared to the data.

Figure 7-5 through Figure 7-8 present the calculated axial void fraction versus the measured data for several points in time during the transient. The comparisons are presented for 10, 40, 100, and 160 seconds into the transient. The results show reasonable-to-excellent agreement. The trend of increased void fraction along the height of the vessel is predicted rather well. Deviations are observed at the lower elevations.

The results of mixture level in the vessel, not presented here, also show reasonable-to- excellent agreement between the calculated NRELAP5 results and the experimental data. The results validate the ability of NRELAP5 to predict key in-vessel thermal- hydraulic phenomena associated with a rapid depressurization event.

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Figure 7-4. GE level swell 1 ft. vessel pressure versus time

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Figure 7-5. GE level swell 1 ft. vessel void fraction versus elevation at 10 seconds

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Figure 7-6. GE level swell 1 ft. vessel void fraction versus elevation at 40 seconds

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Figure 7-7. GE level swell 1 ft. vessel void fraction versus elevation at 100 seconds

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Figure 7-8. GE level swell 1 ft. vessel void fraction versus elevation at 160 seconds

7.2.3 GE Level Swell (4 ft)

7.2.3.1 Facility Description

The 4 ft. GE Level Swell test facility is fully described in Reference 59 and summarized in this section. The experimental facility shown in Figure 7-9 consists of a pressure vessel made of carbon steel with a volume of 160 ft3 (4.5306 m3), 47 in. (1.1938 m) in diameter and 14 ft (4.2672 m) in length. The test facility included a 10 in. (0.254 m) diameter vertical blowdown dip tube to simulate top break locations and a horizontal blowdown line to simulate bottom break locations. The effluent from the vessel blowdown is discharged to a suppression tank.

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Figure 7-9. Schematic of the GE 4 ft. blowdown vessel

Three basic types of measurements were obtained during each experiment: pressures, differential pressures, and temperatures.

As shown in Figure 7-9 the pressure drop is measured at seven sections between the adjacent differential pressure taps, numbered sequentially. Similar to the 1 ft. GE level swell test, the mixture density and void fraction in each measurement section were calculated from the measured pressure drop (see Section 7.2.2.1).

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7.2.3.2 Phenomena Addressed

The phenomena addressed with the 4 ft. GE level swell are same as in the 1 ft. GE level test (see Section 7.2.2.2)

7.2.3.3 Experimental Procedure

The experiment procedure consisted of filling the vessel with demineralized water and boiling the inventory at atmospheric pressure to remove any dissolved gas. The top vent was then closed and the fluid was heated to the starting conditions, which was typically saturated conditions at 1,060 psia (7.308 MPa) for the large blowdown vessel experiments. The initial water level was dependent upon the experiment of interest. Top and bottom break blowdown events were conducted utilizing rupture discs.

A test 5801-15 with top break and initial liquid level of 5.5 ft (1.676 m) is selected for the assessment.

7.2.3.4 Special Analysis Techniques

The {{ }}2(a),(c) improves the depressurization for the sensitivity modeling the discharge and choking into an atmospheric blowdown tank. The discussion here, however, provides a summary with {{ }}2(a),(c)

7.2.3.5 Assessment Results

The results of the GE level swell (4 ft. vessel) from NRELAP5, using the {{ }}2(a),(c) choking model and the measured data are compared. Key parameters are plotted together with the test data in Figure 7-10 through Figure 7-13. The results show reasonable-to-excellent agreement based on the comparison of the pressure and void fractions in the vessel. These results validate the ability of NRELAP5 to predict key in-vessel thermal-hydraulic phenomena associated with a rapid depressurization event.

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Figure 7-10. GE level swell 4-ft vessel pressure versus time

Figure 7-11. GE level swell 4-ft vessel void fraction versus elevation at 5 seconds

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Figure 7-12. GE level swell 4-ft vessel void fraction versus elevation at 10 seconds

Figure 7-13. GE level swell 4-ft vessel void fraction versus elevation at 20 seconds

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7.2.4 KAIST

In the NuScale design, the DHRS is a passive safety-related system that relies on film condensation and natural circulation to remove heat from the RCS through the SG and reject heat to the reactor pool through the DHRS condenser. Reference 60 produced a suitable high pressure steam condensation experimental database which is used to assess the condensation model in NRELAP5.

The KAIST test data varied the pressure and non-condensable gas fraction of the steam entering the test section (mockup of a condenser tube). {{

}}2(a),(c),ECI

7.2.4.1 Facility Description

A schematic of the KAIST test facility is shown in Figure 7-14. Figure 7-15 shows the schematic of the test section. The maximum design pressure and temperature of the test facility were 7.5 MPa (1.088 psia) and 300 degrees C (752 degrees F), respectively.

The major components of the test facility include: SG which supplied steam (maximum power 200 kWe), test section tube, cooling pool (cools the test section), steam line (transports steam from SG to the test section inlet), condensate drain line, LP (or condensate collection tank), and air supply system. The test section was immersed in the cooling pool and was cooled by boiling and single-phase convective heat transfer on the outside surface of the test section.

The test section was a vertical tube with an inside diameter of 4.62 cm (1.82 in.) and an effective heat transfer length of 1.8 m (71 in.). The thickness of the tube wall was 2.3 mm (0.09 in.). To reduce the entrance effect, the top 0.5 m (20 in.) length of the test section was insulated. The test section was submerged in a cooling pool of width 1.2 m × 1.2 m (47 in. x 47 in.) and 2.5 m (98 in.) height. A steam line with an inside diameter 2.34 cm (0.92 in.) was connected from the top of the SG to the top of the test section. The condensate from the test section was drained to the LP (or condensate collection tank) by gravity and then pumped back to the SG.

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Steam Line T 4000 mm P PS SafetyValve Water Atom. CV Pool P

T T T Secondary P T DP Condensor T DP Vent T T T CV Tank T DP T Air Flow Steam DP Generator T CV Air Flow T

T T T Air Supply Air Heat Exchanger Tank Compressor DP Lower T Plenum DP Feed & Drain T

T T CV 200 kW Power Water Line Water Recirculation Electric Heater Water Heat Exchanger Pump

Figure 7-14. Schematic of KAIST test facility

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Figure 7-15. Schematic of the KAIST test section

7.2.4.2 Phenomena Addressed

The phenomena addressed with the KAIST assessment include {{

}}2(a),(c)

7.2.4.3 Experimental Procedure

The experiments were started by purging all non-condensable gas (i.e., air) from the test loop. This was done by supplying steam to the test loop and venting it to the atmosphere through the vent valve located below the test section. After all non-condensable gas was purged, the vent valve was closed and the test section was allowed to fill with the condensate by keeping the condensate drain valve closed. After the test section was completely filled, the SG pressure was increased to the test pressure. As soon as the test pressure was reached, the condensate drain valve was opened and the condensate

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recirculation pump was started. A constant water level in the LP was maintained by control of the recirculation pump flow rate. Data acquisition was started after the process had reached a steady state. Parameter ranges for the KAIST tests are summarized in Table 7-4.

Table 7-4. Range of KAIST test data

Parameter Value Pressure (MPa) 0.794 to 7.457 (115 to 1082 psia) 0.036 to 0.34 Reduced pressure (Pr) (using critical pressure of 220.64 bar (3,200 psia) Inlet steam mass flow (kg/s) 0.01 to 0.1 (0.022 to 0.22 lb/s) Inlet air concentration (percent) 0.0 to 30.0

Prandtl number (Prf) 0.84 to 2.63

Liquid Reynolds number (ReLT) 2,300 to 3,2000

Inlet gas Reynolds number (ReGS) 16,400 to 15,0000

7.2.4.4 Special Analysis Techniques

Based on sensitivity studies, {{

}}2(a),(c)

7.2.4.5 Assessment Results

The results show reasonable-to-excellent agreement between the NRELAP5 calculations and the KAIST measured experimental data, on the comparison of condensed liquid flows, heat transfer coefficients, and inner wall temperatures. This is a result of implementation of the {{ }}2(a),(c) in NRELAP5 (see Section 6.8), which is intended to improve the predicted high pressure condensation response.

Figure 7-16 presents the measured versus calculated heat transfer coefficient for the KAIST steam condensation experiments. The majority of the predictions lie within the experimental uncertainty (28 percent for heat transfer coefficient).

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Figure 7-16. Measured versus predicted heat transfer coefficient

Figure 7-17 through Figure 7-19 present heat transfer coefficient, temperature and mass flow rate versus test section elevation. The majority of the predicted values (all but one) lie within the uncertainty range of the data.

Overall, the results show that NRELAP5 calculations are in excellent agreement with the KAIST measured experimental data. These results validate NRELAP5 for prediction of key thermal-hydraulic phenomena associated {{ }}2(a),(c)

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Figure 7-17. KAIST and NRELAP5 axial heat transfer coefficient

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Figure 7-18. KAIST and NRELAP5 axial inner wall temperature

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Figure 7-19. KAIST and NRELAP5 axial liquid mass flow rate

7.2.5 FRIGG

The FRIGG loop tests for the Marviken boiling heavy water reactor project were executed in four phases by ASEA-ATOM during the years 1967-1970 (Reference 61). These experiments included measurements of axial and radial void distribution, single- phase and two-phase pressure drop, natural circulation mass velocity, stability limits as well as detailed dynamic characteristics, and burnout in natural and forced circulation.

The axial and radial void distribution data as a function of mass flow, inlet sub-cooling, system pressure, and thermal power provide an excellent data set for evaluating the NRELAP5 interphase drag and heat transfer models under two-phase flow conditions. The FRIGG phase 4 (FRIGG-4) tests applied both a non-uniform radial and axial thermal power profile on the heated rod bundle best simulating the power profiles associated with a typical operating reactor core. As such the FRIGG-4 tests are used to assess NRELAP5 performance.

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7.2.5.1 Facility Description

The FRIGG-4 test facility consisted of a vertical circular test section containing 36 electrically heated rods, a riser, a steam separator, a downcomer, a condenser, a pump, and connecting pipes. The power supply for the FRIGG loop was capable of providing a maximum of 8 MW of direct current power to the heated rods in the test section. A schematic of the test loop is shown in Figure 7-20. Figure 7-21 shows the locations of the void and pressure sensors used in the test section. Reference 61 provides detailed information on the characteristics of the facility.

The rod bundle simulated a full-scale boiling heavy water reactor fuel element. Each rod had a 4.365 m (172 in.) heated length and a 13.8 mm (0.543 in.) outside diameter. The bundle also included a 20 mm (0.787 in) outside diameter unheated center rod that supported the prototype reactor core grid spacers. The heated rods were arranged in equal intervals in three rings, the inner ring having six rods, the middle ring twelve rods, and the outer ring eighteen rods. The rod bundles were contained within a 159.5 mm (6.28 in.) ID shroud.

The heated rod bundle had axial and radial thermal power peaking factors typical of an equilibrium reactor core. The FRIGG-4 tests have no thermal power variation in the azimuthal (circumferential) direction. The average heat flux for each test was determined by dividing the total thermal power by the total heated surface. The local heat flux at any given radial or axial zone can be determined by multiplying the measured average heat flux by the radial and axial coordinate scale factors.

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Figure 7-20. FRIGG-4 experimental loop

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Figure 7-21. FRIGG-4 36 rod test section

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Figure 7-22. FRIGG-4 zones for evaluation of radial void distribution

7.2.5.2 Phenomena Addressed

The phenomena addressed with the FRIGG-4 assessment include {{ }}2(a),(c)

Specifically, the FRIGG-4 tests assess the ability of NRELAP5 to predict the void distribution data in a rod bundle geometry as a function of mass flow, inlet sub-cooling, system pressure, and thermal power for evaluating interphase drag and heat transfer models under two-phase flow conditions in the core.

7.2.5.3 Experimental Procedure

Test points were obtained by specifying the core electric power, inlet flow rate, inlet subcooling, and system pressure. Measurements of axial void fractions were collected for each radial zone of the rod bundle.

7.2.5.4 Special Analysis Techniques

There were no special analysis techniques utilized.

7.2.5.5 Assessment Results

One-dimensional NRELAP5 model of the test section was used to analyze this test. Figure 7-23 to Figure 7-26 below show the area-weighted average void fractions in axial zones G1 through G7 for tests 613123, 613130, 613010 and 613118. NRELAP5

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predicted the experimental void fraction data with reasonable agreement justifying use of {{

}}2(a),(c)

Figure 7-23. FRIGG mean void data of NRELAP5 versus Test 613123 data

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Figure 7-24. FRIGG mean void data of NRELAP5 versus Test 613130 data

Figure 7-25. FRIGG mean void data of NRELAP5 versus Test 613010 data

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Figure 7-26. FRIGG mean void data of NRELAP5 versus Test 613118 data

7.2.6 FLECHT-SEASET

The FLECHT-SEASET tests (References 62 and 63) consisted of forced and gravity reflood experiments using electrical heater rods to simulate fuel bundles similar to the Westinghouse 17 x 17 design. The test program was originally designed to study large- break LOCA events. Following the Three Mile Island accident, it was re-oriented to obtain data relevant to small break LOCA events.

Because the NuScale core remains covered with coolant for all design basis LOCA events, reflood phenomena does not occur. However, the test campaign included bundle boil-off tests which are relevant for the NuScale design because the NPM uses boiling in the core to remove heat following a number of accident scenarios that result in actuation of the ECCS. Following ECCS initiation and pressure equalization in the NPM, the RPV and CNV are essentially a pool boiler system with coolant boiled off in the RPV being replaced by an inflow of coolant from the CNV.

7.2.6.1 Facility Description

The facility loop with test section is shown in Figure 7-27. The heater rods were manufactured with a prototypical PWR axial cosine power shape.

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steam boiler P pressure TF TW TF line TW TW P control valve TF DP exhaust 12' TW DP orifice DP TW DP differential pressure cells at DP DP each DP TW TF static pressure foot water TW DP TF from TF TW supply TW fluid thermocouple 0' - 12' steam 0' P wall thermocouple separator TF TW TF TW 0-100 gpm TF 8 turbine meter TF test section 0-18 gpm orifice plate flow meter TF 8 0-6 gpm liquid level turbine rotameter carryover tank meter rotometers Figure 7-27. FLECHT-SEASET experimental facility

7.2.6.2 Experimental Procedure

The FLECHT-SEASET boil-off tests were conducted by filling the 12 ft. tall vessel with approximately 10 ft. of (slightly sub-cooled) water. The power to the heater rods was turned on, and the water was allowed to boil. The test was terminated and reflood initiated when a rod thermocouple registered a temperature greater than or equal to 2,000 degrees F. Three separate boil-off tests are used to assess NRELAP5. These tests were conducted with initial system pressures of 20, 40, and 60 psia.

7.2.6.3 Phenomena Addressed

The phenomenon addressed with the FLECHT-SEASET assessment include {{ }}2(a),(c)

Specifically, the FLECHT-SEASET boil off tests assess the ability of NRELAP5 to predict the axial void profile, mixture level (interfacial drag), and cladding temperature response during boil-off of a PWR core.

7.2.6.4 Special Analysis Techniques

Based on sensitivity studies using one-dimensional components, it is concluded that {{

}}2(a),(c)

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7.2.6.5 Assessment Results

The results for Test 35557 performed at 60 psia are shown in Figure 7-28 through Figure 7-35. The predictions for the void fraction at different elevation are shown in Figure 7-28 through Figure 7-31. The comparisons for the collapsed water levels for all sections are provided in Figure 7-32 through Figure 7-35. While the model and data show reasonable agreement, NRELAP5 over-predicts void fractions as a function of time in most of the core region (Figure 7-28 through Figure 7-31) resulting in a conservative earlier prediction of core uncovery when compared to test data. Similar comparisons were obtained for the test runs at 20 psia and 40 psia.

Figure 7-28. FLECHT-SEASET level 1 void fraction versus time – Test 35557

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Figure 7-29. FLECHT-SEASET level 2 void fraction versus time – Test 35557

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Figure 7-30. FLECHT-SEASET level 3 void fraction versus time – Test 35557

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Figure 7-31. FLECHT-SEASET level 4 void fraction versus time – Test 35557

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Figure 7-32. FLECHT-SEASET level 1 collapsed water level versus time – Test 35557

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Figure 7-33. FLECHT-SEASET level 2 collapsed water level versus time – Test 35557

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Figure 7-34. FLECHT-SEASET level 3 collapsed water level versus time – Test 35557

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Figure 7-35. FLECHT-SEASET level 4 collapsed water level versus time – Test 35557

7.2.7 SemiScale (S-NC-02 and S-NC-10)

The Semiscale test loop modeled a typical PWR. The goal of the Semiscale S-NC-2 and S-NC-10 tests was to obtain experimental data on the natural circulation single-phase and two-phase flow conditions at various system inventories for differing system powers. Three powers were investigated: 30 kW, 60 kW, and 100 kW. At each power level the mass inventory was reduced from at or near 100 percent conditions. With the reduction of primary inventory two-phase flow developed resulting in an enhancement of the total system flow rate. Further reduction in system inventory resulted in a degradation of the total system flow rate.

7.2.7.1 Facility Description

The Semiscale Mod-2A test facility is a full-height 1/1,705 power-to-volume scaled model of a typical four-loop PWR. Only one loop was used for the S-NC-2 and S-NC-10 tests discussed here. The single-loop configuration is shown in Figure 7-36. The reactor coolant pump was removed and replaced by an orifice to model the loss of a seized pump.

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Figure 7-36. Semiscale Mod-2A single (intact) loop test facility configuration

7.2.7.2 Phenomenon Addressed

The phenomenon addressed with the Semiscale assessment includes {{ }}2(a),(c)

Specifically, the Semiscale natural circulation tests assess the ability of NRELAP5 to predict natural circulation during single- and two-phase flow conditions at various system inventories and system powers in a complex geometry.

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7.2.7.3 Experimental Procedure

Prior to initiation of the tests, the primary system was filled with demineralized water and vented to ensure it was liquid-full. The primary system was heated using core power as a heat source and the SG secondary system as a heat sink. Single-phase natural circulation flow driven by density gradients in the loop was used to thermally condition the system to obtain a specified set of initial conditions.

For those steady-state tests in which the primary system was to be drained, the pressurizer was used only to establish initial conditions. The pressurizer was disconnected from the coolant loop prior to draining the primary system.

Primary system mass inventory was controlled by draining fluid from the vessel LP in discrete steps. This fluid was condensed and measured using a static pressure transducer.

The SG secondary levels were controlled by a feed-and-bleed process combined with secondary system draining. The secondary-system pressure was maintained such that saturation conditions prevailed through the use of a steam control valve.

The steady-state natural circulation tests used constant core powers from 30kW to 100kW, representing 1.5 percent to 5.0 percent of the 2,000 kW full Semiscale core power.

During the steady-state experiments, the independent variables were controlled in discrete, step-wise manners, allowing steady conditions to be established between the times when changes in the independent variables were made.

External heaters were used to offset heat losses from the primary coolant system that would affect loop natural circulation behavior. The heaters were located on the hot leg, pump suction, cold leg and vessel downcomer sections of the experiment system. The external heater powers were adjusted to follow previously-determined system heat loss versus system temperature relations. The effectiveness of the external heaters was verified by ensuring constant temperatures (indicative of no heat losses) across these sections.

Three test cases (30kW, 60kW, and 100kW) of S-NC-2 and S-NC-10 were evaluated.

7.2.7.4 Special Analysis Techniques

During the NRELAP5 assessment of the Semiscale test cases, it was noted that the {{

}}2(a),(c)

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7.2.7.5 Assessment Results

The NRELAP5 models for S-NC-02 at 30 kW and 60 kW, and S-NC-10 at 100 kW were run at various system mass inventories. Once steady conditions are obtained for each inventory reduction, the last 200 seconds of the interval are averaged to obtain key FOMs for comparison to data.

The NRELAP5 predictions of loop mass flow rate as a function system inventory are compared to the experimental data in Figure 7-37 to Figure 7-39. In general, NRELAP5 provides reasonably to excellent agreement when predicting the trends, the peak two- phase flow rate, and the enhanced flow rate region (region to the right of the peak). Minor discrepancies are noted to exist in the degraded loop flow region (75 percent – 80 percent inventory levels). These results validate the ability of NRELAP5 to predict natural circulation during single- and two-phase flow conditions at various system inventories and system powers in a complex geometry.

Figure 7-37. S-NC-2 30 kW average mass flow rate versus percent inventory

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Figure 7-38. S-NC-2 60 kW average mass flow rate versus percent inventory

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Figure 7-39. S-NC-10 100 kW average mass flow rate versus percent inventory

7.2.8 Wilson Bubble Rise

The NPM hot leg riser is a large-diameter pipe. During various phases of a LOCA, a nearly stagnant two-phase mixture will be present in the riser. The Wilson bubble rise experimental data are useful to validate NRELAP5 for prediction of void fraction distribution in the hot leg riser.

7.2.8.1 Facility Description

The test facility shown in Figure 7-40 includes a steam inlet and exit nozzle as well as an 18-in. (0.457 m) diameter channel inserted vertically within a 36-in. (0.914 m) diameter vessel. A simplified portion of the test section has been modeled with selected boundary conditions. The boundary conditions consist of the inlet steam mass flow rate and exit pressure. Steam enters the test section from the bottom and exits at the top.

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Figure 7-40. Schematic of Wilson bubble rise test facility

7.2.8.2 Experimental Procedure

The Wilson bubble rise experiments were executed by the vessel being slowly heated and brought to equilibrium at the desired test pressure. The water level and steam flow rates were adjusted to the desired values. After the system reached equilibrium, the necessary instrument readings were taken. These readings were the vessel pressure, steam flow, and the three radial void fraction readings (outer radial region, median region and central region of channel). After the readings were taken, the next steam flow was set and the process was repeated. The steam flow was varied from 5,000 to 60,000 lb/hr (2,268 to 27,216 kg/hr) and the pressures ranged from 600 to 2,000 psi (4.14 to 13.8 MPa).

7.2.8.3 Phenomena Addressed

The phenomenon addressed with the Wilson bubble rise assessment case are {{ }}2(a),(c)

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Specifically, the Wilson bubble rise tests assess the ability of NRELAP5 to predict axial void distribution (dependent on interfacial drag) within a large diameter vertical channel.

7.2.8.4 Special Analysis Techniques

Based on sensitivity studies, {{

}}2(a),(c)

7.2.8.5 Assessment Results

The results in Figure 7-41 through Figure 7-43 show the comparison of predicted and measured void fraction at different pressures. Figure 7-44 shows the data for all cases plotted on predicted versus measured graph. A reasonable agreement is observed between the calculated NRELAP5 results and the Wilson bubble rise measured experimental data. In general NRELAP5 conservatively predicted higher void fraction (or lower mass inventory).

Figure 7-41. NRELAP5 and Wilson void fraction versus superficial velocity at 600 psig (4.14 MPa)

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Figure 7-42. NRELAP5 and Wilson void fraction versus superficial velocity 1,000 psig (6.89 MPa)

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Figure 7-43. NRELAP5 and Wilson void fraction versus superficial velocity 2,000 psig (13.8 MPa)

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Figure 7-44. Predicted versus measured area averaged void fraction (all cases)

7.2.9 Marviken Jet Impingement Test (JIT) 11

The Marviken JIT-11 (Reference 66) was chosen to assess the single-phase choked flow model in NRELAP5.

7.2.9.1 Facility Description

A schematic of the experimental setup is shown in Figure 7-45. The facility consisted of a pressure vessel of fluid at specified conditions, discharge pipe, ball valve, and discharge nozzle. The facility was constructed with focus on measuring loads due to discharged fluid impingement on a flat plate and full-scale critical flow data. The facility was constructed with a stand-pipe such that only single-phase steam was discharged through the break nozzle.

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Figure 7-45. Marviken jet impingement test facility

7.2.9.2 Experimental Procedure

Each test consisted of first obtaining desired initial conditions in the pressure vessel followed by bursting the rupture disk in the discharge pipe. For JIT-11 a stand-pipe was installed in the pressure vessel such that only steam from the upper plenum of the vessel was discharged. The test was conducted at 5 MPa (725 psia) and nearly- saturated liquid in the vessel. The nozzle diameter for was 299.0 mm (0.098 ft) with a nozzle length of 1.18 mm (7.4×10-3 in.).

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7.2.9.3 Phenomenon Addressed

The Marviken JIT-11 addresses the ability of NRELAP5 to predict single-phase (vapor) choked flow (mass and energy release).

7.2.9.4 Special Analysis Techniques

{{

}}2(a),(c)

7.2.9.5 Assessment Results

Figure 7-46 and Figure 7-47 compare the experimental data and the NRELAP5 simulated mass flow rate and density for various values of the discharge coefficient. Excellent agreement is shown with the experimental data for {{

}}2(a),(c)

{{

}}2(a),(b),(c) Figure 7-46. Marviken jet impingement test 11 flowrate

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{{

2(a),(b),(c) }} Figure 7-47. Marviken jet impingement test 11 density

7.2.10 Bankoff Perforated Plate

Bankoff, et al. (References 67, 68, and 69) conducted air/water and steam/water counter current flow tests in a small scale test apparatus that established counter current flow through a number of different perforated plates. The Bankoff correlation assessment uses the CCFL implementation as described in Section 6.6.3.

7.2.10.1 Facility Description

Reference 6 9 describes the Bankoff CCFL test apparatus. Additional information on the test apparatus and additional tests are reported in References 67 and 68. A horizontal perforated plate is located in a vertical test assembly. Steam or air can be introduced below the plate and water can be injected above the plate. A water overflow line is located above the plate to limit the height of the “bubbly pool” of water above the plate. The perforated plate could be moved so that the height of the “bubbly pool” could be varied. There is a drain at the bottom of the test section to prevent water level from building up below the plate. A beam scale is placed at the drain to measure the flow of water that penetrates through the plate. Air or steam that was not condensed on the injected water exited at the top of the test apparatus. The test simulated for this assessment used a 15-hole plate with a “bubbly pool” height of 267

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mm (10.5 in.). The test was conducted at atmospheric pressure. A schematic of the test facility is shown in Figure 7-48.

Figure 7-48. Schematic of Bankoff counter current flow apparatus (from Reference 68)

7.2.10.2 Phenomenon Addressed

The phenomenon addressed with the Bankoff assessment case is CCFL at pressurizer baffle plate and upper core plate (UCP) (or top nozzle).

7.2.10.3 Experimental Procedure

The test was conducted by establishing a water inlet flow rate and then increasing the air flow rate in a stepwise manner. The rate of water flow through the perforated plate was measured by weighing the flow out of the bottom of the test section. The test was concluded when the air flow was sufficient to prevent water downflow through the perforated plate.

7.2.10.4 Assessment Results

Figure 7-49 compares predicted vapor superficial velocity versus predicted liquid superficial velocity. The comparison shows that the predictions are in excellent agreement with the experimental data.

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Figure 7-49. Superficial vapor velocity versus superficial liquid velocity

7.2.11 Marviken Critical Flow Test 22 and 24

The Marviken critical flow tests (CFTs) (Reference 70 and 71) were conducted to characterize the conditions of blowdown given {{ }}2(a),(b),(c)

7.2.11.1 Facility Description

{{

}}2(a),(b),(c)

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{{

}}2(a),(b),(c) Figure 7-50. Schematic of the Marviken pressure vessel

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{{

}}2(a),(b),(c) Figure 7-51. Discharge pipe dimensions and instrument locations

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7.2.11.2 Experimental Procedure

{{

}}2(a),(b),(c)

7.2.11.3 Phenomenon Addressed

The phenomenon addressed with the Marviken assessment is two-phase and single- phase choked flow.

7.2.11.4 Special Analysis Techniques

{{

}}2(a),(b),(c)

7.2.11.5 Assessment Results

7.2.11.5.1 Comparison to Marviken Critical Flow Test-22

{{

}}2(a),(b),(c)

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{{

}}2(a),(b),(c)

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{{

}}2(a),(c)

Figure 7-52. Measured versus calculated mass flow rate for Marviken critical flow test 22

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{{

}}2(a),(c)

Figure 7-53. Marviken critical flow test 22 comparison to calculated mixture density

7.2.11.5.2 Comparison to Marviken Critical Flow Test-24

{{

}}2(a),(b),(c)

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{{ }}2(a),(b),(c)

{{

}}2(a),(c)

Figure 7-54. Measured versus calculated mass flow rate for Marviken critical flow test 24

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{{

}}2(a),(c)

Figure 7-55. Marviken critical flow test 24 mixture density and calculated mixture density

Analysis shows that NRELAP5 has the capability to perform critical flow calculations with reasonable-to-excellent agreement to test data.

7.3 NuScale Stern Critical Heat Flux Tests

The CHF correlations described in Section 6.11 were assessed against steady state CHF experiments performed by NuScale in the Stern facility, and this assessment is presented in this section.

The Stern tests were performed on a preliminary prototypical bundle geometrically comparable to the NuFuel HTP2™ design, but with {{ }}2(a),(b),(c),ECI The Stern preliminary prototypical bundle tests provide data over wide parameter ranges, which encompass the NPM operating parameter values and can be used to assess the capability of NRELAP5 to predict the onset of the CHF. Key FOMs to assess agreement include the critical power and the critical power ratio as a function of mass flux, pressure, and inlet sub-cooling.

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7.3.1 Facility Description

The Stern CHF tests made use of a 5x5 fuel bundle comprised of {{ }}2(a),(b),(c),ECI heated length fuel simulators arranged in three configurations including:

• {{ }}2(a),(b),(c) (U-1 series) • {{ }}2(a),(b),(c) (U-2 series) • {{ }}2(a),(b),(c) (C-1 series).

A prototypical fuel diameter {{

}}2(a),(b),(c),ECI

{{

}} 2(a),(b),(c),ECI

Figure 7-56. U1 & C1 (left) versus U2 (right) radial layout

An axial layout of the test section with key instrument locations is shown in Figure 7-57. The test section includes a pressure housing, a channel box (flow channel), fuel simulators, spacer grids, and instrumentation. Four spacers are installed within the heated section of the assembly at prototypical locations with a spacer pitch of {{ }}2(a),(b),(c),ECI The resistance temperature detectors are used to measure the average inlet and outlet temperatures of the coolant. In addition to the absolute pressure measurements at the inlet and outlet of the test section, there

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are nine differential pressure transducers installed within the heated section to measure the pressure drop across various axial sections.

{{

}}2(a),(b),(c),ECI

Figure 7-57. Stern test section axial layout

7.3.2 Experimental Procedure

At the Stern test facility, the steady state CHF tests were performed in the following manner:

• loop conditions were established with the heated assembly at a power below the critical power,

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• loop conditions were maintained steady as much as possible while the power was {{ }}2(a),(b),(c) until the critical power was reached, • the data acquisition program continuously scans the assembly signals and critical power is considered to occur when the {{

}}2(a),(b),(c),ECI • when the occurrence of critical power is confirmed the loop conditions were held steady and the steady-state data was recorded, and • once the test point was recorded the power was reduced, as necessary, and loop conditions changed for the next test.

7.3.3 Phenomenon Addressed

The Stern CHF benchmark assesses the ability of NRELAP5 to predict CHF and {{ }}2(a),(c)

7.3.4 Parameter Ranges Assessed

The Stern steady state CHF tests were conducted across a systematic range of mass flows, inlet pressures, and inlet sub-cooling. A total of {{ }}2(a),(b),(c),ECI steady state CHF data were collected for pressures ranging from {{

}}2(a),(b),(c),ECI as described in Table 7-5. A series of repeat tests was also performed to determine the repeatability of the test data. A total of {{ }}2(a),(b),(c),ECI repeat test points are identified. Only the {{ }}2(a),(b),(c),ECI high flow data points with mass fluxes greater than {{ }}2(a),(b),(c),ECI are excluded from the assessment presented in this document.

Table 7-5. Range of Stern steady state critical heat flux data {{

}}2(a),(b),(c),ECI

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7.3.5 Special Analysis Techniques

The NRELAP5 model consists of a {{

}}2(a),(c)

7.3.6 Assessment Results

{{

}}2(a),(c)

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{{

}}2(a),(b),(c),ECI Figure 7-58. Predicted versus measured Stern power

7.4 NuScale SIET Steam Generator Tests

This section addresses assessments performed against experiments conducted under the NuScale testing program at SIET laboratories, in Piacenza, Italy. Two test programs were conducted as described below.

7.4.1 SIET Tests

The SIET experimental program is a two test activity with helical coil SG tubes characterized on an electrically heated test section (TF-1), and on a fluid heated test section (TF-2). The electrically heated test provides detailed in-tube information for the secondary side, while the fluid heated test allows investigation of the general behavior of the tube bundle heated by the primary side fluid. This section deals with the detailed description of the electrically heated test (TF-1) and NRELAP5 assessment results.

The electrically heated test section incorporates three full scale coils of the helical coil SG, providing information focused on the SG secondary side. Direct heating of the test section is provided by passing current through the tubes using three different axial heating zones (subcooled, saturated and superheat).

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{{ }}2(a),(b),(c),ECI coils of the electrically heated test section represent the {{ }}2(a),(b),(c),ECI coils of the NuScale SG, in terms of diameter, length and angle of inclination, and they allow investigation into the effects of tube curvature on thermal-hydraulic parameters. The {{ }}2(a),(b),(c),ECI coil reproduces the {{ }}2(a),(b),(c),ECI coil of the NuScale SG, in terms of diameter and length, {{

}}2(a),(b),(c),ECI

7.4.1.1 Facility Description

The main components and loops of the SIET TF-1 facility in the NuScale helical coil SG test configuration are described here. A pump system drives water from a water storage tank to the pre-heating zone where it is brought to the specified operating conditions and sent to a feedwater header. The header feeds the three coils of the test section that can be activated by valves: singularly or two in parallel. Superheated steam exits the test section toward a header connected to the separation and discharge system. A schematic of the test loop is provided in Figure 7-59.

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{{

}}2(a),(b),(c),ECI Figure 7-59. SIET electrically-heated test instrumentation diagram

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7.4.1.2 Phenomena Addressed

The SIET TF-1 assessment cases addresses {{ }}2(a),(c)

7.4.1.3 Experimental Procedure

For adiabatic tests the inlet flowrate is specified along with the outlet pressure for each test point. For diabatic tests the inlet temperature, flowrate, and tube/zone heat flux (by setting the current) are specified along with the outlet pressure. This section only covers diabatic tests.

{{

}}2(a),(b),(c)

7.4.1.4 Special Analysis Techniques

The helical coil component used includes the helical coil friction model and heat transfer packages inside the coil (see Section 6.7).

7.4.1.5 Assessment Results

In general, NRELAP5 predicted the experimental data with reasonable-to-excellent agreement. The following specific conclusions were drawn from the assessment

• Calculated axial fluid and wall temperatures are within reasonable-to-excellent agreement of data. • Calculated single- and two-phase pressure drops along the coil are in reasonable-to- excellent agreement with the test data.

Results from two Coil 2 tests are presented first to illustrate variation between predicted and measured wall temperature along the length of the coil. Subsequently, pressure drop for tests on Coil 1 through 3, and fluid temperature and wall temperature for tests on Coil1 are presented.

Wall temperature profile for the three heating zones (subcooled, saturated, and superheat) of coil 2 are depicted in Figure 7-60 and Figure 7-61 for diabatic tests TD0015 and TD0003, respectively. From inspection of the wall temperatures, {{

}}2(a),(b),(c),ECI

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{{

}}2(a),(b),(c),ECI

{{

}}2(a),(b),(c),ECI

Figure 7-60. Time averaged wall temperature profile for coil 2 test TD0015

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{{

}}2(a),(b),(c),ECI

Figure 7-61. Time averaged wall temperature profile for coil 2 test TD0003

Pressure drops for the five sections along the length of Coil 1 (i.e., axial pressure drop) are given in Figure 7-62. The error bands on these figures represent the uncertainty in measurement of pressure drops. Calculated pressure drops over the first {{ }}2(a),(b),(c),ECI are predicted with excellent agreement and within the experimental error, as shown by Figure 7-62. Similar results are shown for pressure drops in Coil 2 and Coil 3 in Figure 7-63 and Figure 7-64. In general, NRELAP5 does a reasonable-to-excellent job of predicting the axial pressure drops taking into account that the standard deviation of the experimental data (not shown on plots) is larger than the reported measurement uncertainty (shown on plot).

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{{

}}2(a),(b),(c),ECI

Figure 7-62. SIET electrically-heated test differential pressure for all coil 1 diabatic tests

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{{

}}2(a),(b),(c),ECI

Figure 7-63. SIET electrically-heated test differential pressure for all coil 2 diabatic tests

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{{

}}2(a),(b),(c),ECI

Figure 7-64. SIET electrically-heated test differential pressure for all coil 3 diabatic tests

Fluid temperatures for {{ }}2(a),(b),(c),ECI of coil 1 are depicted in Figure 7-65. The error bands on these figures represent the uncertainty in measurement of fluid temperature. The calculated values are in reasonable-to-excellent agreement with experimental data.

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{{

}}2(a),(b),(c),ECI

Figure 7-65. SIET electrically-heated test fluid temperatures for all coil 1 diabatic tests

Corresponding wall temperatures at several axial locations of coil 1 are depicted in Figure 7-66. The error bands on these figures represent the uncertainty in measurement of wall temperature. Wall temperature results are similar to the corresponding fluid temperature results {{

}}2(a),(b),(c),ECI

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{{

}}2(a),(b),(c),ECI

Figure 7-66. SIET electrically-heated test wall temperature for all coil 1 diabatic tests

7.4.2 SIET Fluid-Heated Test

The SIET fluid-heated tests were performed in support of the NuScale design development, with particular emphasis on providing experimental data for validation of NRELAP5 for prediction of helical coil SG primary and secondary heat transfer, primary side pressure drop, and secondary side dryout.

7.4.2.1 Facility Description

The SIET TF-2 facility consists of a 252 helical coil tube bundle installed inside a pressure vessel. The tube bundle consists of 5 tube banks, simulating the {{

}}2(a),(b),(c),ECI All five tube bundles are placed in an annulus, formed by two cylindrical barrels, installed axially within the pressure vessel. The helical coils are wrapped around the inner barrel and kept in position by four supports, {{ }}2(a),(b),(c),ECI

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{{

}}2(a),(b),(c),ECI

Each tube bank is fed by a feed-water vertical header, inside the vessel, that distributes water to each helical tube. Steam from each the exit of each tube bank is collected in a steam vertical header and driven outside the vessel top nozzle by pipes.

The primary-side of the test section consists of an inlet riser barrel, connection bellows, internal barrel, pressure vessel dome, free volume (i.e., unoccupied by the tubes) between the internal and external barrels (i.e., annulus), and free volume of the pressure vessel around the inlet riser and connection bellows. Water on the primary side is circulated by pumps and pre-heated by an electric heater before entering the pressure vessel. The pressure on the primary side is maintained using the electrically heated pressurizer.

Primary water, entering the pressure vessel from the bottom nozzle, rises through a vertical channel and enters the central cylindrical part of the test section, representing the riser. After reaching the vessel dome, water turns down into the test section annulus to cross the helical coil tube bundle. Exiting at the bottom, water is driven again to the circulation pumps.

Instruments are installed for the measurement of primary side mass flow rate, inlet and exit temperatures, pressures and differential pressures. Instruments are installed to measure the secondary side feed water flow rate, feed water temperature, pressures and differential pressure along the tubes, and exit steam temperatures and flow rates.

7.4.2.2 Phenomenon Addressed

Adiabatic tests were performed to characterize the primary side pressure losses in the facility. These tests were run without heat input to the primary flow and there was no secondary flow to the coils.

Diabatic tests measured pressure drop and heat transfer on both the primary and secondary, and the thermal crisis (dryout) location on the secondary side during heated operation of the coils. These tests characterize the thermal performance of the coils for a range of primary and secondary side inlet flows and temperatures.

7.4.2.3 Experimental Procedure

Target boundary conditions are obtained for the diabatic tests. The duration of the data recording for each test was a minimum 300 seconds for the pre-steady state and 150 seconds for the steady state.

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7.4.2.4 Special Analysis Techniques

This benchmark assesses {{

}}2(a),(c)

7.4.2.5 Assessment Results

In general, NRELAP5 predicted the experimental data with reasonable-to-excellent agreement. The following specific conclusions were drawn from the assessment:

• {{

}}2(a),(b),(c),ECI

7.4.2.5.1 Assessment of Adiabatic Experiment Data

Adiabatic experimental data from TF-2 testing is used to assess the modeling of primary side friction and form losses. The primary side pressure drop was measured at {{ }}2(a),(b),(c),ECI Figure 7-67 shows the comparison of predicted and measured primary side pressure drop at all axial elevations for all adiabatic tests. The error bands represent the uncertainty in measurement of pressure drops. Excellent agreement between NRELAP5 predictions and measured test data exists with primary side pressure drop predicted within the measurement uncertainty. Similar results are obtained for other primary side pressure drop measurement elevations. {{

}}2(a),(b),(c),ECI

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{{

}}2(a),(b),(c),ECI Figure 7-67. SIET fluid-heated test adiabatic primary differential pressure

7.4.2.5.2 Primary Side Pressure Drop and Fluid Temperatures of Diabatic Experiments

{{

}}2(a),(b),(c),ECI

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{{

}}2(a),(b),(c),ECI Figure 7-68. SIET fluid-heated test diabatic test primary differential pressure

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{{

}}2(a),(b),(c),ECI

Figure 7-69. SIET fluid-heated test diabatic test primary temperature

7.4.2.5.3 Steam Generator Tube Wall Temperature of Diabatic Experiments

{{

}}2(a),(b),(c),ECI

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}}2(a),(b),(c),ECI Figure 7-70. Comparison of wall temperatures in TD0001 (Case 1A)

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}}2(a),(b),(c),ECI Figure 7-71. Comparison of wall temperatures in TD0005 (Case 1A)

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}}2(a),(b),(c),ECI Figure 7-72. Comparison of wall temperatures in TD0015 (Case 1A)

7.4.2.5.4 Secondary Side Fluid Temperature of Diabatic Experiments

Figure 7-73 and Figure 7-74 show the comparison of predicted and measured secondary side fluid temperatures at all elevations in Row 3 for selected tests. The figures also show the predicted and measured primary fluid temperatures. {{

}}2(a),(b),(c),ECI

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{{

}}2(a),(b),(c),ECI

Figure 7-73. Comparison of primary and secondary side fluid temperatures in TD0001 (Case 1A)

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{{

}}2(a),(b),(c),ECI

Figure 7-74. Comparison of primary and secondary side fluid temperatures in TD0005 (Case 1A)

7.5 NuScale NIST-1 Test Assessment Cases

A scaled facility of the NPM was constructed at Oregon State University, referred to as the NIST-1 facility, to assist in validation of the NRELAP5 system thermal-hydraulic code. The facility is designed to perform various tests, including LOCA tests. The NIST-1 facility consists of the major components in the NPM. These components include: an RPV, helical coil SG system with DHRS, CNV, and cooling pool vessel (CPV) representing the reactor pool. The NIST-1 ECCS connects the RPV to the CNV and consists of two RVVs and two RRVs, each on separate lines. Breaks can be simulated for the RCS lines that connect the RPV to the CNV to simulate piping breaks within the CNV. This system consists of a RCS discharge line, a RCS injection line, and a pressurizer spray supply line. The CVCS is not functional in the NIST-1 facility and is used only for simulation of CVCS line break LOCAs.

Instrumentation is included in the facility to capture the response of the system under steady-state and transient situations. The instrumentation includes pressure, differential pressure, water level, mass flow rate, heat flux, and temperature measurements.

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7.5.1 Test Facility Description

Due to the unique nature of the NPM design the number of IET facilities suitable for code assessment is limited. What is now the NIST-1 facility was originally conceived at OSU in 2000 as a proof-of-concept testing platform for development of Small Modular Reactor (SMR) technology. During this period it was referred to as the multi-application small light water reactor facility (Reference 72).

Although the NuScale design was based on the Multi-Application Small Light Water Reactor (MASLWR), the concept has evolved considerably since the inception of NuScale in 2008. At the time that NuScale was formed, the facility was renamed the NIST facility. The NIST facility is a scaled, non-nuclear reactor that uses electric heater rods to represent the heat produced from fission. It is designed to produce experimental data in support of verification and validation of thermal-hydraulic codes.

In 2014 and 2015, the original NIST facility was modified by NuScale to facilitate accurate simulation and to bring the facility in-line with the current NuScale plant design configuration. Following the upgrade, the NIST facility was renamed NIST-1 facility. A scaling analysis was employed for design of the NIST test facility to ensure that the facility design is capable of capturing important plant phenomena with minimal distortions. Further discussion on the NIST-1 facility scaling and distortions is available in Sections 8.3.2 and 8.3.4.

Updates to the NIST facility included in NIST-1 are:

• {{

}}2(a),(b),(c),ECI

The updated NIST-1 facility provides a well-scaled representation of the current NuScale reactor design that minimizes distortions and provides the measurements necessary for safety code and reactor design validation. A schematic of the NIST-1 facility is shown in Figure 7-75.

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{{

}}2(a),(c),ECI

Figure 7-75. Schematic of NuScale integral test facility and NRELAP5 nodalization

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The NIST-1 facility models the NPM at {{ }}2(a),(b),(c),ECI scale. There are three vessels in the NIST-1 facility: the RPV, CNV, and CPV as shown in Figure 7-75. Unlike the plant, the RPV and CNV are not concentric and the CNV is not immersed in the CPV. Rather the RPV and CNV are connected by piping that contains valves that perform the functions of the RRVs, RVVs and breaks as shown in Figure 7-75. This approach enables flow measurements to be made in this piping during testing. The CNV is connected to the CPV through a HTP that is scaled to allow energy transfer to the pool in the same proportion as in the NPM.

Natural circulation flow in the primary circuit is driven by heat input in the core region and heat removal to the SG tubes. Fluid heated in the core region flows upward through the hot leg riser, and then downward around the outside of the SG tubes, the cold leg and the downcomer. The flow then returns to the core through the LP. The core is comprised of a {{ }}2(a),(b),(c),ECI electric heater rod bundle with a maximum power of {{ }}2(a),(b),(c),ECI kW, a power level scaled to simulate decay heat. System pressure is controlled by the pressurizer component which contains heater rods to bring the pressurizer fluid up to saturation temperature.

7.5.1.1 Reactor Pressure Vessel

Major internal components in the RPV are the core, hot leg riser, pressurizer, and SG bundle. The pressurizer is at the top, separated from the lower part of the RPV by a perforated pressurizer baffle plate. The upper plenum occupies the region below the pressurizer baffle plate and above the hot leg riser that extends down to the top of the core. The upper annulus between the hot leg riser and the RPV shell contains the helical coil SG tubes. The lower part of the annulus immediately below the SG tubes is the cold leg. The lower annulus at the core elevation is the downcomer, which is separated from the core by the core shell. The LP occupies the bottom of the RPV and hydraulically connects the downcomer and the core.

The RPV shells and flanges are covered by {{

}}2(a),(b),(c),ECI

7.5.1.1.1 Reactor Core

The RPV houses the core, which is modeled by a {{

}}2(a),(b),(c),ECI

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{{ }}2(a),(b),(c),ECI

7.5.1.1.2 Hot Leg Riser

After leaving the core, the flow enters the chimney of the hot leg riser. The hot leg riser extends from above the core shroud to the upper plenum, creating a riser and downcomer configuration to enable natural circulation. The hot leg riser consists of a lower shell, a conical transition, a middle shell containing the flowmeter for the primary circuit, and an upper shell. Flow exits the riser into the upper plenum, which is the space between the hot leg riser outlet and the bottom of the pressurizer baffle plate.

7.5.1.1.3 Upper Plenum

After leaving the top of the hot leg riser, the flow enters the upper plenum and is directed radially outward to flow down in the annulus between the riser and the RPV shell. The pressurizer baffle plate separates the upper plenum from the pressurizer. Hydraulic communication between the pressurizer and the RPV occurs via holes located in the pressurizer baffle plate.

7.5.1.1.4 Pressurizer

The pressurizer is located above the upper plenum and is in thermal-hydraulic communication with the upper plenum via the pressurizer baffle plate holes. The pressurizer maintains primary system pressure during normal steady-state and transient conditions through the use of three heater elements. Each element has {{ }}2(a),(b),(c),ECI of power and is modulated by the facility control system to maintain system pressure.

7.5.1.1.5 Cold Leg Downcomer

After leaving the upper plenum, the flow continues downward through the SG section and into the cold leg downcomer region. The cold leg downcomer is the annular space bounded by the RPV shell ID and the hot leg riser outer diameter. When fluid reaches the hot leg riser conical transition shell, the flow area is reduced. Flow exits the cold leg downcomer into the LP before it recirculates back into the core.

7.5.1.1.6 Steam Generators

The SG is a helical-coil, once-through heat exchanger consisting of {{ }}2(a),(b),(c),ECI that wrap counter to each other in the annular space between the hot leg riser and the RPV shell inner surface. In the NIST-1 facility, the primary coolant is circulated on the outside of the SG tubes, similar to the NPM. Feedwater supplied from the feedwater storage tank is pumped through the SG coils by a regenerative turbine pump. Pressure in the secondary side is regulated by a pneumatically operated variable position valve located in the steam line portion of the flow loop.

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7.5.1.1.7 Lower Plenum

The LP is the region bounded between the tubesheet and the lower core flow plate. The LP provides the connection between the downcomer and the core, thus completing the RPV flow loop.

7.5.1.2 Containment

The CNV, representing the cavity volume between the RPV outer surface and the containment inside surface, is conjoined to the CPV and thermally separated by a scaled HTP. For the NPM, the RPV is located inside containment. However, with the NIST-1 facility, to maintain both volume and surface area scaling similitude, as well as allow proper instrumentation, the RPV is thermal hydraulically separated from the CNV. The CNV models the scaled condensation heat transfer surface between the CNV and CPV. Fluid in the CPV, which is at ambient pressures, models the scaled volume in which an NPM CNV is submerged.

7.5.1.3 Cooling Pool Vessel

The CPV has a set of four ports allowing for the installation of one of three decay heat removal heat exchangers. The baseline configuration is with a {{

}}2(a),(b),(c)

7.5.1.4 Emergency Core Cooling System and Chemical and Volume Control System Lines

Eight lines connect the RPV to the CNV. Five of these lines belong to the facility ECCS, whereas the other three are part of the CVCS. As part of the ECCS, there are two independent reactor vent lines near the top of the pressurizer section and two reactor recirculation lines in the lower downcomer of the RPV. The fifth ECCS line is an SET line that also models reactor recirculation. For the CVCS, two lines penetrate the vessel near the bottom of the SG. One of these lines penetrates both the vessel wall and the hot leg riser, simulating the make-up line into the hot leg. The other CVCS line connects to the cold leg and penetrates only the RPV wall. This line represents the facility CVCS discharge break line. A third CVCS line between the RPV and CNV is located at the top of the pressurizer and functions as an analogy for the pressurizer spray supply line. Each line has a pneumatic isolation valve that is actuated through the test facility control system. Any lines that are not installed use a blank flange for isolation.

7.5.1.5 Facility Instrumentation and Control

Instrumentation is used throughout the NIST-1 facility to measure the thermal-hydraulic behavior during steady-state and transient operations. The following information is recorded by the DACS:

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• {{

}}2(a),(b),(c),ECI The data generated and collected by the facility DACS is used to validate applicability of the NRELAP5 thermal-hydraulic code for LOCA analysis.

7.5.1.6 Integral Effects LOCA Test Procedure

Prior to startup, a valve and switch lineup is performed to place the facility in the desired configuration for the upcoming test. The break line modeling the break location specified for the test is connected between the RPV and its associated CNV penetration. To prevent an accidental actuation of an incorrect break valve a blind flange is installed in all other break lines. Orifices with the specified diameters are installed in the RVV and RRV lines to model the number of valves that are to open when ECCS actuates.

Because the NIST-1 facility has a nominal operating pressure of {{ }}2(a),(b),(c),ECI that is less than the NPM pressure of 1,850 psia (12.76 MPa), the test in the NIST-1 facility simulates the NPM transient in progress. Specifically, the RPV and CNV fluid masses in NIST-1 are scaled such that they are {{ }}2(a),(b),(c),ECI that of the RPV and CNV fluid masses in the NPM at a corresponding pressure of {{ }}2(a),(b),(c),ECI Thus the initial RCS mass inventory and pressure are preserved on a scaled basis and fluid property similitude is maintained throughout the transient.

As part of the NIST-1 LOCA tests, {{

}}2(a),(b),(c),ECI

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{{

}}2(a),(b),(c),ECI

7.5.2 Facility NRELAP5 Model

The NRELAP5 model of the NIST-1 facility is constructed to {{

}}2(a),(c) These model features are shown in the NRELAP5 nodalization shown in Figure 7-75.

7.5.3 Facility Test Matrix

This NIST-1 facility is used to perform design certification IETs and SETs for the purpose of validating NuScale computer codes, model development and assessment, correlation development, verifying compliance with design requirements, demonstrating design features and capabilities, and addressing regulatory concerns.

This section briefly describes the test matrix for the NIST-1 facility. Descriptions of tests used for NRELAP5 code validation are provided in Table 7-6. These are the NIST-1 tests that are the essential subset of tests required to validate NRELAP5 for NPM LOCA calculations.

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Table 7-6. Facility high priority tests for NRELAP5 code validation {{

}}2(a),(b),(c),ECI

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{{

}}2(a),(b),(c),ECI Tests NIST-1 HP-06, HP-07 and HP-09 are the IETs that are used for validating the NRELAP5 EM for LOCA applications. Test HP-09 is included because spurious opening of an RVV results in the bounding RPV depressurization rate. Tests NIST-1 HP-43 and HP-49 were performed to support the extension of LOCA EM for the analysis of transients initiated due to inadvertent opening of RPV valves. Further discussion on the NRELAP5 validation results against these tests is provided in Appendices B and C. These tests also supported the containment response analysis methodology.

7.5.4 Separate Effect High Pressure Condensation Tests

The NIST-1 facility HP-02 test was used to assess the capability of NRELAP5 to predict condensation rates at high pressure test conditions by comparing experimental data and NRELAP5 predictions.

7.5.4.1 Facility Description

The HP-02 test is an SET performed at the NIST-1 facility. The test involves injecting steam at known conditions into the CNV and measuring the CNV pressure, level, and temperature response. Only the CNV, CPV, and interconnecting HTP are important to this test. During testing, the RPV was pressurized and heated using core heat to supply superheated steam from the SG to the CNV at the desired mass flow rate.

The feedwater flowrate was measured with individual Coriolis flowmeters to each of the three SG inlet tube banks. Also, one Coriolis meter measured the total SG feedwater inlet flow and one vortex flowmeter measured the total steam flow at the SG exit. The Coriolis flowmeter measuring the combined inlet flow was used as a mass flow boundary

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condition in the NRELAP5 model as it provided the most stable flow measurement with the lowest measurement uncertainty.

7.5.4.2 Phenomenon Addressed

The pertinent phenomena validated with the NIST-1 facility HP-02 assessment are {{

}}2(a),(b),(c),ECI

7.5.4.3 Experimental Procedure

Initial steam conditions in the CNV were obtained by first operating the NIST-1 facility in its normal mode, heating the RPV with core heaters with heat rejection through the SG to the environment. The SG feedwater flowrate, core power, and steam exit pressure were established to obtain the desired conditions for steam. Once the desired conditions were established, steam was diverted from the stack (rejected to the environment) to the CNV.

Five tests were run to evaluate steady-state condensation at varying CNV pressures. For each test, superheated steam was discharged into the CNV until the CNV target pressure was reached, after which the inlet steam flow was ramped down in an effort to achieve steady state conditions at the target pressure.

Steady steam inlet conditions were maintained through the injection period. After steam was injected into the CNV, condensation occurred on the HTP. Condensation energy was then thermally conducted through the HTP and convected into the CPV.

7.5.4.4 Parameter Ranges Assessed

Test conditions were selected to obtain condensation data at various CNV pressures. Five tests were conducted at steady CNV pressure varying from {{

}}2(a),(b),(c),ECI

7.5.4.5 Assessment Results

The HP-02 test facility data was compared to NRELAP5 predictions designed to simulate the test conditions and test procedures in effect during the experiment. HP-02 test data trends were well predicted by NRELAP5 with reasonable-to-excellent agreement for condensation rates at pressures ranging {{ }}2(a),(b),(c),ECI

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{{ }}2(a),(b),(c),ECI NRELAP5 has demonstrated its capability to predict CNV level, CNV pressure, CNV temperature, and CPV temperature with reasonable-to-excellent agreement.

The following subsections provide a brief summary of the results for three HP-02 runs analyzed.

7.5.4.5.1 Run 1 Results

Both the CNV pressure and level responses for Run 1 depicted in Figure 7-76 and Figure 7-77 are in reasonable-to-excellent agreement with the data. The NRELAP5- simulated pressure peak occurs at the same time as the data; reaching a maximum of {{

}}2(a),(b),(c)

The CNV and CPV fluid temperatures predicted by NRELAP5 are in excellent agreement with the data. Figure 7-78 and Figure 7-79 show that the predictions closely following the data trend and magnitude during the earlier transient as well as the steady- state period.

TR-0516-49422-NP-A Rev. 2 {{

}}2(a),(b),(c),ECI

Figure 7-76. HP-02 Run 1 containment vessel pressure response

TR-0516-49422-NP-A Rev. 2 {{

}}2(a),(b),(c),ECI

Figure 7-77. HP-02 Run 1 containment vessel collapsed level response

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{{

}}2(a),(b),(c),ECI

Figure 7-78. HP-02 Run 1 upper containment vessel fluid temperature response (in vapor space)

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{{

}}2(a),(b),(c),ECI

Figure 7-79. HP-02 Run 1 upper cooling pool vessel temperature response

7.5.4.5.2 Run 2 Results

During run 2 a maximum pressure of {{ }}2(a),(b),(c),ECI was reached. NRELAP5 is in reasonable-to-excellent agreement with the experimental data for CNV pressure, as shown in Figure 7-80. NRELAP5 predicts the general trends for level, with reasonable-to-excellent agreement to data (Figure 7-81), but slightly underpredicts collapsed level.

The NRELAP5 containment vessel and CPV temperatures shown in Figure 7-82 and Figure 7-83 are in excellent agreement with the data, closely following the trend and lying almost entirely within the instrument uncertainty.

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{{

}}2(a),(b),(c),ECI

Figure 7-80. HP-02 Run 2 containment vessel pressure response

TR-0516-49422-NP-A Rev. 2 {{

}}2(a),(b),(c),ECI

Figure 7-81. HP-02 Run 2 containment vessel collapsed level response

TR-0516-49422-NP-A Rev. 2 {{

}}2(a),(b),(c),ECI

Figure 7-82. HP-02 Run 2 upper containment vessel fluid temperature response (in vapor space)

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{{

}}2(a),(b),(c),ECI

Figure 7-83. HP-02 Run 2 upper cooling pool temperature response

7.5.4.5.3 Run 3 Results

{{

}}2(a),(b),(c),ECI

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{{

}}2(a),(b),(c),ECI

Figure 7-84. HP-02 Run 3 containment vessel pressure response

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{{

}}2(a),(b),(c),ECI

Figure 7-85. HP-02 Run 3 containment vessel collapsed level response

TR-0516-49422-NP-A Rev. 2 {{

}}2(a),(b),(c),ECI

Figure 7-86. HP-02 Run 3 upper containment vessel fluid temperature response (in vapor space)

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{{

}}2(a),(b),(c),ECI

Figure 7-87. HP-02 Run 3 upper cooling pool temperature response

Based on this assessment, NRELAP5 has demonstrated its capability to predict CNV level, CNV pressure, CNV temperature, and CPV temperature with reasonable-to- excellent agreement for high pressure condensation conditions.

7.5.5 Natural Circulation Test at Power

The NIST-1 test HP-05 was used to assess the capability of NRELAP5 to predict natural circulation flow at various core powers and test conditions by comparing experimental data and NRELAP5 predictions.

7.5.5.1 Facility Description

The HP-05 test configuration uses the RPV and SG to drive steady-state natural circulation within the RPV at various core heater rod power levels. Core heater rods supply energy to heat the working fluid which, due to buoyancy forces, travels up the riser entering the upper plenum. The fluid then turns 180 degrees and passes over the integrated helical coil SG, exchanging energy to the secondary side. The primary working fluid exits the SG traveling downward through the downcomer, entering the LP

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where another 180 degree turn (upward) is made into the entrance of the electrically heated core.

Various instruments measure differential pressures, flow, temperatures, pressures, and heater power to assess the loop flowrate and pressure losses.

7.5.5.2 Phenomenon Addressed

The pertinent phenomena addressed with the HP-05 assessment case are {{

}}2(a),(c)

7.5.5.3 Experimental Procedure

The HP-05 experiment consists of inducing a core power ramp at a constant RPV pressure of approximately {{ }}2(a),(b),(c),ECI and a secondary-side pressure of approximately {{ }}2(a),(b),(c),ECI Differential pressures around the primary loop were measured to characterize the pressure drops due to form and friction losses. The mass flow rate in the riser and fluid temperatures around the loop are measured. To facilitate comparing to code predictions the core power and temperature rise across the core are used to calculate a theoretical flowrate based on an energy balance.

Test HP-05 initiates from a power of {{ }}2(a),(b),(c),ECI, at a pressure of {{ }}2(a),(b),(c),ECI , and the steady-state natural circulation flow condition. Once steady-state conditions are achieved, {{

}}2(a),(b),(c),ECI

7.5.5.4 Special Analysis Techniques

{{

}}2(a),(c) The global response was then confirmed by comparing the experimental loop flow rate to that predicted by NRELAP5.

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7.5.5.5 Parameter Ranges Assessed

{{

}}2(a),(b),(c)

7.5.5.6 Assessment Results

{{ }}2(a),(b),(c) The NRELAP5 mass flow signal is taken from the same location. The NRELAP5 prediction is closely aligned with the data and shows excellent agreement, with the exception of the behavior demonstrated at the lowest core power level, where reasonable agreement is obtained. At the lower power level, facility constraints on the secondary side made it difficult to obtain steady state conditions.

{{

}}2(a),(c) {{

}}2(a),(b),(c),ECI

Figure 7-88. HP-05 NIST-1 averaged mass flowrate and NRELAP5 results

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The core inlet temperature was measured in the LP upstream of where the fluid enters the core. The NRELAP5 signal is taken from the same region. Comparisons to the measure data are provided in Figure 7-89. The NRELAP5 core inlet temperatures are in reasonable agreement with the data.

{{

}}2(a),(b),(c),ECI

Figure 7-89. HP-05 NIST-1 averaged core inlet temperature and NRELAP5 results

Core outlet temperature was measured in the riser near the core exit. The NRELAP5 signal is located in the same region. Each time the core power is lowered the hot leg temperature first falls and then recovers. In the data, the temperature usually over- shoots the previous steady state value prior to settling down at the next steady state value. Except for these power transients, the data and NRELAP5 predictions are in excellent agreement as demonstrated in Figure 7-90 except the oscillations observed in NRELAP5 at low power/flow conditions.

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{{

}}2(a),(b),(c),ECI

Figure 7-90. HP-05 NIST-1 averaged core outlet temperature and NRELAP5 results

Based on this assessment, NRELAP5 has demonstrated its capability to predict primary flow rate, core inlet temperature, and core outlet temperature with reasonable-to- excellent agreement for natural circulation flow conditions.

7.5.6 Chemical and Volume Control System Loss-of-Coolant Accident Integral Effects Tests

The HP-06 test was used to assess the capability of NRELAP5 to predict the integral response of the NIST-1 facility for a single-ended discharge line break inside containment. The discharge line and valve connect the downcomer side of the RPV to the CNV.

The HP-06b test was similar to the HP-06 test, with the exception of the core power. This test was performed to assess the impact of core power on the progression of the LOCA. {{

}}2(a),(b),(c),ECI

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7.5.6.1 Facility Description

The entire NIST-1 facility except for the DHRS was used for this IET, including:

• the SG was active to remove heat from the primary side and drive natural circulation in conjunction with the electrically heated core during the steady state • {{

}}2(a),(b),(c) • the CPV was filled to accept rejected heat from the HTP In addition, {{

}}2(a),(b),(c)

7.5.6.2 Phenomenon Addressed

The HP-06 and HP-06b tests are IETs modeling a single-ended discharge line break inside containment. The purpose of these IETs was to assess the integral response of the scaled NIST-1 facility. The pertinent phenomena addressed by these tests are:

• {{

}}2(a),(b),(c),ECI

7.5.6.3 Experimental Procedure

The IET test procedure is described in Section 7.5.1.6. When the CNV pressure reached the specified CNV break initiation pressure, the CVCS break valve was opened, initiating the transient.

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short period of time. The CNV pressurization and heat-up occurs rapidly, followed by a long depressurization and cooldown profile. Test data was recorded for an extended period of time, well into the long-term cooling phase.

7.5.6.4 Special Analysis Techniques

The RCS discharge line orifice has a length of approximately {{ }}2(a),(b),(c),ECI and an ID of approximately {{ }}2(a),(b),(c),ECI Thus, the orifice has an L/D ratio roughly equal to {{ }}2(a),(b),(c),ECI Analysis indicates that a NRELAP5 discharge coefficient near {{ }}2(a),(b),(c),ECI produces reasonable agreement with the break flow test data.

The {{

}}2(a),(c)

7.5.6.5 Assessment Results (HP-06)

The NRELAP5 transient model is designed to simulate initial test conditions and includes logic that follows facility controls and test procedures. The NRELAP5-calculated RCS discharge line break mass flow rate is shown in Figure 7-91 with a peak flowrate of approximately {{ }}2(a),(b),(c),ECI lbm/s. For this experiment, the break mass flow rate was not measured. The calculated break flow rate is reasonable because the differential pressure across the RCS discharge line orifice (Figure 7-92), the RPV level response (Figure 7-95), the CNV level response (Figure 7-96), the RPV pressure response (Figure 7-99), and the CNV pressure response (Figure 7-97) are all in excellent agreement.

The NIST-1 v-cone flowmeter (measuring primary loop flowrate) is designed for positive single-phase liquid conditions. During the HP-06 test, two-phase conditions occur at the location of the v-cone meter. Figure 7-93 shows that NRELAP5 predicts the RPV primary-flow coast-down after break initiation with reasonable accuracy. The measured RPV mass flow rate after approximately 29 seconds post-test initiation is more uncertain due to potential for two-phase conditions at the v-cone meter.

The pressurizer level is compared in Figure 7-94. The comparisons show reasonable-to- excellent agreement. NRELAP5 predicts complete draining of the pressurizer at about {{ }}2(a),(b),(c),ECI

NRELAP5 provides reasonable-to-excellent agreement for level response in the RPV and CNV as shown in Figure 7-95 and Figure 7-96. The CNV peak pressure and pressure response are also predicted with excellent agreement to data as shown in Figure 7-97 and Figure 7-98. The timing of ECCS actuation is predicted with reasonable- to-excellent agreement to the test data. Primary pressure response is predicted with reasonable-to-excellent agreement (Figure 7-99).

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{{

}}2(a),(b),(c),ECI

Figure 7-91. NIST-1 HP-06 NRELAP5 chemical and volume control system discharge line break mass flow rate

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{{

}}2(a),(b),(c),ECI

Figure 7-92. NIST-1 HP-06 break orifice differential pressure

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{{

}}2(a),(b),(c),ECI

Figure 7-93. NIST-1 HP-06 primary mass flow rate

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{{

}}2(a),(b),(c),ECI

Figure 7-94. NIST-1 HP-06 pressurizer level comparison

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{{

}}2(a),(b),(c),ECI

Figure 7-95. NIST-1 HP-06 reactor pressure vessel level comparison

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{{

}}2(a),(b),(c),ECI

Figure 7-96. NIST-1 HP-06 containment vessel level comparison

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{{

}}2(a),(b)(c),ECI

Figure 7-97. NIST-1 HP-06 containment vessel pressure comparison

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{{

}}2(a),(b),(c),ECI

Figure 7-98. NIST-1 HP-06 containment vessel pressure comparison {{

}}2(a),(b),(c),ECI

Figure 7-99. NIST-1 HP-06 primary pressure comparison

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7.5.6.6 Assessment Results (HP-06b)

{{

}}2(a),(b),(c),ECI

Other HP-06b initial and boundary conditions were similar to the HP-06 test {{

}}2(a),(b),(c),ECI

{{

}}2(a),(b),(c),ECI

Figure 7-100. Comparison of core power in HP-06 and HP-06b tests with the NuScale Power Module decay power after reactor trip (scaled)

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Figure 7-101 and Figure 7-102 show the comparisons of predicted and measured RPV and CNV pressures, respectively. Similar comparisons for the RPV and CNV levels are shown in Figure 7-103 and Figure 7-104, respectively. Overall, NRELAP5 predicted the HP-06b data with reasonable-to-excellent agreement.

Figure 7-105 and Figure 7-106 show the comparisons of measured RPV and CNV pressures in HP-06 and HP-06b tests, respectively. Similar comparisons for the measured RPV and CNV levels are shown in Figure 7-107 and Figure 7-108, respectively. {{

}}2(a),(b),(c),ECI {{

}}2(a),(b),(c),ECI

Figure 7-101. NIST-1 HP-06b primary pressure comparison

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{{

}}2(a),(b),(c),ECI

Figure 7-102. NIST-1 HP-06b containment vessel pressure comparison

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{{

}}2(a),(b),(c),ECI

Figure 7-103. NIST-1 HP-06b reactor pressure vessel level comparison

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{{

}}2(a),(b),(c),ECI

Figure 7-104. NIST-1 HP-06b containment vessel level comparison

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{{

}}2(a),(b),(c),ECI

Figure 7-105. Comparison of NIST-1 HP-06 and HP-06b reactor pressure vessel pressure

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{{

}}2(a),(b),(c),ECI

Figure 7-106. Comparison of NIST-1 HP-06 and HP-06b containment vessel pressure

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{{

}}2(a),(b),(c),ECI

Figure 7-107. Comparison of NIST-1 HP-06 and HP-06b reactor pressure vessel level

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{{

}}2(a),(b),(c),ECI

Figure 7-108. Comparison of NIST-1 HP-06 and HP-06b containment vessel level

7.5.7 Pressurizer Spray Supply Line Loss-of-Coolant Accident Integral Effects Test

The HP-07 test was used to assess the capability of NRELAP5 to predict the integral response of the NIST-1 facility modeling a single-ended pressurizer spray supply line break inside containment.

7.5.7.1 Facility Description

The entire NIST-1 facility, except for the DHRS, was used for this IET.

• The SG was active to remove heat from the primary side and drive natural circulation in conjunction with the electrically heated core during steady state. • {{

}}2(a),(b),(c) • The CPV was filled to accept rejected heat from the HTP.

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7.5.7.2 Phenomenon Addressed

The phenomena addressed in the test facility HP-07 test are same as in the HP-06 test (see Section 7.5.6.2).

7.5.7.3 Experimental Procedure

The LOCA test procedure is discussed in Section 7.5.1.6. When the CNV pressure reached the specified CNV break initiation pressure, the pressurizer spray supply line break valve was opened, initiating the transient.

Within the NIST-1 facility, the ECCS actuation occurs when the compensated level in the RPV downcomer reaches a pre-specified value. Once this occurs, open signals are sent to the RRVs and the RVVs. The opening of the ECCS valves causes a large amount of mass and energy transfer to occur between the RPV and the CNV over a short period of time. Containment vessel pressurization and heat-up occurs rapidly, followed by a long depressurization and cooldown profile. Test data was recorded for an extended period of time, well into the long-term cooling phase.

7.5.7.4 Special Analysis Techniques

The pressurizer spray supply line orifice has a length of approximately {{ }}2(a),(b),(c),ECI. Thus, the orifice has an L/D ratio roughly equal to {{ }}2(a),(b),(c),ECI. Analysis indicates that a NRELAP5 discharge coefficient near {{ }}2(a),(b),(c),ECI produces the best overall agreement with the break flow test data.

The {{

}}2(a),(c)

7.5.7.5 Assessment Results

Figure 7-109 shows the comparison of core power in the HP-07 test to the scaled NPM total power after reactor trip (i.e., fission product decay, actinide decay, and fission power). The HP-07 power is approximately representative {{ }}2(a),(b),(c),ECI of the NPM power.

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{{

}}2(a),(b),(c),ECI

Figure 7-109. Comparison of core power in HP-07 with the NuScale Power Module power (fission and decay) after reactor trip (scaled)

The break flowrate predicted by NRELAP5 (Figure 7-110) provided results with excellent agreement to data with the use of the {{ }}2(a),(c) choking model and a discharge coefficient of {{ }}2(a),(b),(c),ECI

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}}2(a),(b),(c),ECI

Figure 7-110. NIST-1 HP-07 pressurizer spray supply line break discharge mass flow rate

{{

}}2(a),(b),(c),ECI

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{{

}}2(a),(b),(c),ECI

Figure 7-111. NIST-1 HP-07 primary mass flow rate

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{{

}}2(a),(b),(c),ECI

Figure 7-112. NIST-1 HP-07 reactor pressure vessel level response comparison with data

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{{

}}2(a),(b),(c),ECI

Figure 7-113. NIST-1 HP-07 containment vessel level response

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{{

}}2(a),(b),(c),ECI

Figure 7-114. NIST-1 HP-07 containment vessel pressure comparison

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{{

2(a),(b),(c),ECI }} Figure 7-115. NIST-1 HP-07 primary pressure comparison

7.5.8 Spurious Reactor Vent Valve Opening Test

The HP-09 test was used to assess the capability of NRELAP5 to predict the integral response of the NIST-1 facility to inadvertent depressurization of the RPV initiated by RVV spurious opening without DHRS. Furthermore, this test also provided bounding depressurization rate for a LOCA initiated by break from pressurizer gas space.

7.5.8.1 Facility Description

The entirety of the NIST-1 facility except for the DHRS was used for this IET.

• The SG was active to remove heat from the primary side and drive natural circulation in conjunction with the electrically heated core during the steady state. • {{

}}2(a),(b),(c)

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• {{

}}2(a),(b),(c) • The CPV was filled to accept rejected heat from the HTP.

7.5.8.2 Phenomenon Addressed

The phenomena addressed in the NIST-1 HP-09 test are same as in the HP-06 and HP- 07 IETs (see Section 7.5.6.2).

7.5.8.3 Experimental Procedure

The experimental procedure is consistent with the LOCA test procedure discussed in Section 7.5.1.6.

7.5.8.4 Special Analysis Techniques

The {{ }}2(a),(c) at the valve orifice. Furthermore, the modified PV term was activated at the valve orifice. The Bankoff CCFL model was applied at pressurizer baffle plate.

Analysis indicates that a NRELAP5 discharge coefficient near {{ }}2(a),(b),(c),ECI produces the best overall agreement with the valve flow test data.

7.5.8.5 Assessment Results

Figure 7-116 shows the comparison of core power in the HP-09 test to the scaled NPM total power after reactor trip (i.e., fission product decay, actinide decay, and fission power). The HP-09 core power is {{ }}2(a),(b),(c),ECI of the power in NPM.

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{{

}}2(a),(b),(c),ECI

Figure 7-116. Comparison of HP-09 core power with the scaled NuScale Power Module fission and decay power

Figure 7-117 through Figure 7-124 compare NIST-1 HP09 test data with the NRELAP5 transient response. The calculated RVV mass flow rate is shown in Figure 7-117 with a peak flowrate of approximately {{ }}2(a),(b),(c),ECI lbm/s. The mass flow rate is over- predicted during the first {{ }}2(a),(b),(c),ECI seconds of the transient. Thereafter, the calculated flow shows excellent agreement with the measured flow rate.

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{{

}}2(a),(b),(c),ECI

Figure 7-117. NIST-1 HP-09 valve mass flow rate

The calculated pressurizer pressure is compared to data in Figure 7-118. The calculated pressure shows excellent agreement with the data, including the time of ECCS initiation. An examination of the first {{ }}2(a),(b),(c),ECI seconds of the RPV pressure (Figure 7-119) shows that the NRELAP5 predicted pressure is higher than the measured pressure.

Figure 7-120 compares the NRELAP5-predicted and NIST-1-measured CNV pressure response. Figure 7-121 shows the short-term response. The peak pressure from data and model are {{ }}2(a),(b),(c),ECI psia and {{ }}2(a),(b),(c),ECI psia, respectively. The comparison shows reasonable-to-excellent agreement with the measured data. As with the RPV pressure response, after about {{ }}2(a),(b),(c),ECI seconds, the CNV pressure is under-predicted. The trends of the data are well represented in the calculation.

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{{

}}2(a),(b),(c),ECI

Figure 7-118. NIST-1 HP-09 pressurizer pressure comparison

{{

}}2(a),(b),(c),ECI

Figure 7-119. NIST-1 HP-09 pressurizer pressure comparison

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{{

}}2(a),(b),(c),ECI

Figure 7-120. NIST-1 HP-09 containment vessel pressure comparison

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{{

}}2(a),(b),(c),ECI

Figure 7-121. NIST-1 HP-09 containment vessel pressure comparison

The pressurizer and RPV levels are compared in Figure 7-122 and Figure 7-123, respectively. The comparisons show reasonable-to-excellent agreement. Note that the code-to-data comparison presented in Figure 7-122 shows that NIST-1 pressurizer draining fully occurs between {{ }}2(a),(b),(c),ECI seconds, i.e., when the data (LDP-1401_calc) reaches a value of approximately {{ }}2(a),(b),(c),ECI inches, the low range of the measurement. When the data is extrapolated out past {{ }}2(a),(b),(c),ECI, it appears full draining of the pressurizer occurs at about {{ }}2(a),(b),(c),ECI seconds. NRELAP5 predicts pressurizer draining to fully occur at about {{ }}2(a),(b),(c),ECI seconds.

NRELAP5 predicts {{ }}2(a),(b),(c),ECI As shown in Figure 7-123 the RPV level prediction is in reasonable-to-excellent agreement with the test data.

A closer look at the RPV level comparison over the first {{ }}2(a),(b),(c),ECI seconds (Figure 7-124) shows excellent agreement.

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{{

}}2(a),(b),(c),ECI

Figure 7-122. NIST-1 HP-09 pressurizer level comparison

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{{

}}2(a),(b),(c),ECI

Figure 7-123. NIST-1 HP-09 reactor pressure vessel level comparison

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{{

}}2(a),(b),(c),ECI

Figure 7-124. NIST-1 HP-09 reactor pressure vessel level comparison

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8.0 Assessment of Evaluation Model Adequacy

The adequacy of the NRELAP5 code (Reference 9) for analysis of design-basis LOCAs in the NPM is demonstrated by closure model and correlation reviews, and assessments against relevant experimental data. Establishing the adequacy of the NRELAP5 code as a component of the NuScale LOCA methodology is an essential part of the EMDAP (RG 1.203).

8.1 Adequacy Demonstration Overview

The adequacy demonstration process used here is similar to that used for the AP-600 (Reference 74). As the NuScale PIRT is a primary input to the adequacy evaluation, the findings of the PIRT summarized in Section 4 are used to guide the adequacy of the evaluation process. The adequacy of the NuScale LOCA EM is demonstrated through the following steps:

1. Section 8.2 documents the bottom-up assessment of the NRELAP5 models and correlations to determine their adequacy to predict the high (H) ranked phenomena. The code models used to represent each high (H) ranked phenomena are identified, with emphasis on the phenomena with low-knowledge level. These assessments address the fidelity of the models and correlations to the appropriate fundamental or SET data. Fidelity of the assessments is evaluated using the criteria of excellent, reasonable, minimal and insufficient from RG 1.203. These criteria are defined in Table 1-2. The comparisons to data can identify modeling deficiencies which could impose limitations on the application of the NRELAP5 based LOCA EM. 2. Section 8.3 covers the top-down assessment of the EM including a review of EM governing equations and numerics to determine their applicability to NPM LOCA analysis, and evaluation of the integral code performance based on the assessments of the EM against relevant IETs. 3. Section 8.4 summarizes the adequacy findings. The report shows how each PIRT high (H) ranked phenomenon is covered by the LOCA methodology models and correlations. Models which are marginally adequate, or ranges where PIRT phenomena are not covered, are identified. The manner of addressing code limitations is described.

8.2 Evaluation of Models and Correlations (Bottom-Up Assessment)

The adequacy of the models and correlations in NRELAP5 for modeling the high (H) ranked phenomena is examined by considering their pedigree, applicability, and fidelity to appropriate fundamental or SET data (established by assessment of the EM against legacy and NuScale-specific SET data), and scalability to the NuScale LOCA scenario.

The following steps are used to perform the evaluations.

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{{

}}2(a),(c)

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1. {{

}}2(a),(c)

The result of these assessments for each model or correlation is used to identify whether there are any shortcomings in the parametric space and provide information needed for the development effort where additional models, assumptions, or conservatisms may be required.

The first three steps described above are addressed in Section 8.2.1. Step 4 is discussed in Sections 8.2.2 through 8.2.22.

8.2.1 Important Models and Correlations

Table 8-1 identifies the dominant code models and correlations for the PIRT, defined as high-ranked phenomena in Section 4. Key parameters that are influenced by the dominant models and correlations are listed, along with phenomenological and separate effects cases that are used to assess the model or correlation capabilities. This information is used to establish adequacy of the dominant code models or correlations for NPM LOCA applications.

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Table 8-1. Dominant NRELAP5 models and correlations

{{

}}2(a),(c)

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{{

}}2(a),(c)

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{{

}}2(a),(c) Table 8-2 is a summary of the estimated range of key parameters over which each dominant model or correlation should be applicable for the NPM steady-state and design basis LOCA. Parameter ranges obtained are intended to identify the minimum range that needs to be covered; the applicability of models and correlations are not restricted to these ranges. Several sources are used to obtain the values of the ranges. This includes design values, proposed technical specification limits, and limiting initial and boundary conditions. The ranges for some parameters are obtained from NRELAP5 LOCA break spectrum calculations described in Section 9.0. An explanation of how each limiting range was determined is provided in the Comments column of Table 8-2.

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Table 8-2. NuScale Power Module range of process parameters

{{

2(a),(c),ECI }}

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{{

}}2(a),(c),ECI Table 8-3 lists the range of geometric parameters that could influence high-ranked phenomena for the NPM LOCA. Values given for each parameter are intended to identify the minimum range geometric parameters that need to be covered by the LOCA EM; applicability of the NuScale LOCA EM is not restricted to these values. These values have been obtained from a compilation of geometric information and plant parameters determined from design drawings.

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Table 8-3. Range of NuScale Power Module geometric parameters {{

}}2(a),(c),ECI

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}}2(a),(c),ECI

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Each of the NRELAP5 dominant models or correlations listed in Table 8-1 has been evaluated with respect to the extent that the model or correlation, as assessed for the NPM LOCA application, covers all or a portion of the NPM range given in Table 8-2. Where the range provided in the model or correlation development does not cover the full range of the NPM LOCA application, the range is extended by extrapolation of assessments against experimental data, or justification is provided based on legacy RELAP5-3D© assessments and applications. The range covered by models and correlations is discussed for the key parameters that define the response for high-ranked phenomena.

8.2.1.1 Applicability Evaluation

To determine adequacy of the models and correlations to simulate the high-ranked phenomena, the results of assessments against phenomenological and SETs are discussed. The assessment results are drawn from the NRELAP5 assessments discussed in Section 7.0 where descriptions of the test facilities, instrumentation, and test procedures are provided.

8.2.1.2 Overview

A graded approach is used to address the items described in Section 8.2, Step 4. More emphasis is given to high-ranked phenomena with a low-knowledge level. Less emphasis is placed on phenomena that are well understood with a high-knowledge level. This includes industry standard and handbook models.

Each of the following four areas is evaluated to the extent that they are relevant for each high-ranked phenomenon.

{{

}}2(a),(c)

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8.2.1.2.1 High-Ranked, Low-Knowledge Level Phenomena

The PIRT identified some phenomena within specified components as high-importance phenomena that have a low-knowledge level. These high-importance and low-knowledge phenomena are given the greatest focus in the development of the LOCA EM. They include:

{{

}}2(a),(c)

8.2.2 {{ }}2(a),(c)

8.2.2.1 Background

{{

}}2(a),(c)

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}}2(a),(c)

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}}2(a),(c)

8.2.2.2 Technical Evaluation

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}}2(a),(c)

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Table 8-4. Marviken range of parameters compared to the NuScale Power Module {{

}}2(a),(c),ECI

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}}2(a),(c)

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}}2(a),(c)

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}}2(a),(c)

8.2.3 {{ }}2(a),(c)

8.2.3.1 Background

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}}2(a),(c)

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}}2(a),(c)

8.2.3.2 Technical Evaluation

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}}2(a),(c)

Table 8-5. Ferrell-McGee range of parameters compared to the NuScale Power Module {{

}}2(a),(c),ECI

{{

}}2(a),(c)

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}}2(a),(c)

8.2.4 {{ }}2(a),(c)

8.2.4.1 Background

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}}2(a),(c)

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}}2(a),(c)

8.2.4.2 Technical Evaluation

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}}2(a),(c)

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Table 8-6. Dimensions of NuScale Power Module, NIST-1 and Bankoff pressurizer plate {{

}}2(a),(c),ECI

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}}2(a),(c)

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}}2(a),(c)

8.2.5 {{ }}2(a),(c)

8.2.5.1 Background

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}}2(a),(c)

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}}2(a),(c)

8.2.5.2 Technical Evaluation

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}}2(a),(c)

8.2.6 {{ }}2(a),(c)

8.2.6.1 Background

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}}2(a),(c)

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}}2(a),(c)

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}}2(a),(c)

8.2.6.2 Technical Evaluation

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}}2(a),(c)

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}}2(a),(c)

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Table 8-7. Range of riser interphase friction - separate effects tests and NuScale Power Module {{

}}2(a),(c),ECI {{

}}2(a),(c)

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}}2(a),(c)

8.2.7 {{ }}2(a),(c)

8.2.7.1 Background

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}}2(a),(c)

8.2.7.2 Technical Evaluation

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}}2(a),(c)

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}}2(a),(c)

8.2.8 {{ }}2(a),(c)

8.2.8.1 Background

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}}2(a),(c)

8.2.8.2 Technical Evaluation

Figure 8-1 shows a schematic of the major heat transfer modes governing the heat transfer across the CNV wall. {{

}}2(a),(c)

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}}2(a),(c)

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}}2(a),(c)

Figure 8-1. CNV wall heat transfer modes

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}}2(a),(c)

Figure 8-2. Thermal resistance network between CNV and UHS

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}}2(a),(c)

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}}2(a),(c)

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}}2(a),(c)

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}}2(a),(c)

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8.2.9 {{ }}2(a),(c)

8.2.9.1 Background

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}}2(a),(c)

8.2.9.2 Technical Evaluation

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}}2(a),(c)

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8.2.10 Flashing

8.2.10.1 Background

Flashing is the fundamental thermodynamic process of vaporization that occurs when a saturated liquid undergoes a reduction in pressure below its boiling point, resulting in a phase change from liquid to vapor. In the pressurizer, the liquid inventory that is normally at saturated conditions will flash as the RCS depressurizes in response to a LOCA and the actuation of the ECCS. As the RPV continues to depressurize, flashing will also occur in the hot leg riser, core, LP, and downcomer. Flashing will cause level swell which can affect the quality at the break and at the ECCS valves.

The interphase heat and mass transfer models in NRELAP5 are the dominant models that determine the flashing rate. The vapor generation (or condensation) consists of two parts, vapor generation which results from energy exchange in the bulk fluid (flashing) and energy exchange in the thermal boundary layer near the wall (boiling). Flashing is addressed in this section and boiling in Section 8.2.19.

Each of the vapor generation processes involves interfacial heat transfer effects. The interfacial heat transfer area and heat transfer coefficient models used in NRELAP5 are summarized in Table 2.5-1 of Reference 9. The models that govern the flashing phenomenon are those for the superheated liquid fluid state. For bubbly flow bulk interfacial heat transfer between the vapor and liquid phases is handled using the maximum of a correlation derived from the Plesset-Zwick (Reference 90) equation for the growth rate of a bubble and the modified Lee-Ryley (Reference 91) correlation. For droplets in flow regimes such as annular mist and dispersed, NRELAP5 uses a heat transfer coefficient kf/D f(ΔTsf) where kf is the liquid thermal conductivity, D is hydraulic diameter, and ΔTsf is the difference between the saturation and the liquid temperatures; f(ΔTsf) is a flow-type dependent function. A {{ }}2(a),(c) heat transfer coefficient is used for films. In all cases, a large heat transfer coefficient is calculated so that the difference between the superheated liquid temperature and the saturation temperature at any time is small. Hence for the NPM, while the pressure is decreasing, the liquid temperature remains very close to the saturation temperature due to the relatively slow depressurization resulting from the small break sizes. The time constant for the vapor generation process is much smaller than the time constant for the depressurization. A non-equilibrium superheated liquid state can exist for a time period on the order of milliseconds while the depressurization process is taking place over a period of minutes. Hence, high accuracy in the vapor generation rate due to flashing is not required because any model with a small time constant will generate the amount of vapor necessary to keep the phases in thermal equilibrium.

Implementation of these correlations is described in Reference 9 (Section 2.5.1.1). The pedigree of the model is established by application and validation of RELAP5-3D©. In Table 2.2-2 of Reference 87, it is noted that the flashing model is validated against the

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Edward pipe and Marviken CFT-22 and CFT-24 tests. Validation of NRELAP5 for the suite of assessment cases confirms that this pedigree is maintained in the NRELAP5 code.

8.2.10.2 Technical Evaluation

Important parameters associated with flashing phenomenon are pressure, void fraction/interfacial area, and phasic temperatures that interact to determine the vapor generation rate. The flashing model in NRELAP5 covers the entire range of the water properties tables, which encompasses the NPM LOCA application.

The GE level swell (1-ft and 4-ft) tests are the primary assessment cases used to validate the applicability of NRELAP5 to predict the flashing phenomenon. In Section 8.2.6, it is shown that the NRELAP5 predictions of the GE level swell assessments show reasonable-to-excellent agreement with the test data, thus demonstrating that NRELAP5 is applicable for predicting flashing phenomenon that occurs in NPM LOCA events. Furthermore, flashing is an inherent phenomenon in the NIST-1 LOCA IETs. Therefore, NRELAP5 analysis against the NIST-1 IET data provides additional assessment of the flashing model.

Because the heat transfer coefficient resulting from the Plesset-Zwick correlation depends only on fluid properties and all of the heat transfer coefficients are very large, there are no scaling restrictions for the vapor generation model in NRELAP5 which could impose limitations on the application of the NRELAP5 model to the configuration and conditions of the NPM hot leg riser in the LOCA transient domain.

8.2.11 {{ }}2(a),(c)

8.2.11.1 Background

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}}2(a),(c)

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8.2.11.2 Technical Evaluation

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Figure 8-3. Transient void fraction in node 5 for the GE 4-ft level swell test

Figure 8-4. Transient void fraction in node 4 for the GE 4-ft level swell test

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Figure 8-5. Transient void fraction in node 6 for the GE 1-ft level swell test

Figure 8-6. Transient void fraction in node 5 for the GE 1-ft level swell test

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}}2(a),(c)

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8.2.12 {{ }}2(a),(c)

8.2.12.1 Background

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8.2.12.2 Technical Evaluation

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}}2(a),(c)

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8.2.13 {{ }}2(a),(c)

8.2.13.1 Background

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}}2(a),(c)

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8.2.13.2 Technical Evaluation

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8.2.14 {{ }}2(a),(c)

8.2.14.1 Background

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}}2(a),(c)

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}}2(a),(c)

8.2.14.2 Technical Evaluation

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}}2(a),(c) {{ }}2(a),(c),ECI {{ }}2(a),(c) {{ }}2(a),(c),ECI {{

}}2(a),(c)

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8.2.15 {{ }}2(a),(c)

8.2.15.1 Background

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8.2.15.2 Technical Evaluation

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}}2(a),(c)

8.2.16 {{ }}2(a),(c)

8.2.16.1 Background

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8.2.16.2 Technical Evaluation

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8.2.17 {{ }}2(a),(c)

8.2.17.1 Background

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}}2(a),(c)

8.2.17.2 Technical Evaluation

Table 8-1 lists the flow, void fraction, pressure, heat rate, and core geometry as key parameters for the NRELAP5 interfacial drag model, the dominant NRELAP5 model that affects phase slip and flow regimes. Table 8-2 identifies the range of parameters encountered by NPM steady-state and design basis accidents and transients. This is the source of the ranges used for the NPM {{ }}2(a),(c) SETs used to validate the NRELAP5 interphase drag model for the core geometry are FRIGG tests 613130, 613010, 613118 and 613123, and FLECHT-SEASET boil off tests 35557, 35658 and 35759

The FRIGG tests are steady-state runs with set inlet and boundary conditions to a 36- rod electrically-heated bundle. The void fraction profile along the heated channel was measured and compared to NRELAP5 predictions (see Section 7.2.5).

The FLECHT-SEASET tests (see Section 7.2.6) are boil off transient tests in which a 161-rod bundle was initially filled with saturated water and then boiled to the point at which heater rod temperature reached 2,000 degrees F (1,093 degrees C), and the test was terminated by cutting the rod power and flooding the bundle.

Table 8-8 shows the ranges of the key variables for the NRELAP5 {{

}}2(a),(c)

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Table 8-8. Ranges of key parameters for core interphase friction - separate effects tests and plant {{

}}2(a),(c),ECI {{

}}2(a),(c)

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}}2(a),(c)

8.2.18 {{ }}2(a),(c)

8.2.18.1 Background

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8.2.18.2 Technical Evaluation

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}}2(a),(c)

Table 8-9. Range of key parameters for core flow – separate effects tests and plant {{

}}2(a),(c),ECI

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}}2(a),(c)

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}}2(a),(c)

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8.2.19.1 Background

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8.2.19.2 Technical Evaluation

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Table 8-10. Range of key parameters for core boiling - separate effects tests and plant {{

}}2(a),(c) {{

}}2(a),(c)

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}}2(a),(c)

8.2.20 {{ }}2(a),(c)

8.2.20.1 Background

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}}2(a),(c).

8.2.20.2 Technical Evaluation

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8.2.21 {{ }}2(a),(c)

8.2.21.1 Background

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8.2.21.2 Technical Evaluation

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Table 8-11. Range of key parameters for subcooling boiling and separate effects tests and plant {{

}}2(a),(c) {{

}}2(a),(c)

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8.2.22.1 Background

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8.2.22.2 Technical Evaluation

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8.3 Evaluation of Integral Performance (Top-Down Assessment)

There are three primary areas addressed by the top-down assessment.

• {{

}}2(a),(c)

To ensure maximum fidelity of the assessments, the NRELAP5 NIST-1 and NPM input models were developed using consistent nodalization and option selection. Code assessments are also performed against SETs to establish code capabilities for predicting local behavior within unique NPM components. Assessments against SETs are addressed in Section 8.2.1.

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8.3.1 Review of Code Governing Equations and Numerics

The NRELAP5 Theory Manual (Reference 9) describes the NRELAP5 code architecture, field equations, and solution techniques, which are essentially unchanged compared to the RELAP5-3D© code. The descriptions of code modifications/features made to address unique aspects of the NuScale application are included in the NRELAP5 Theory Manual and summarized in Section 6.0. This review is based primarily on the information in Reference 9.

The field equations solved by NRELAP5 are discussed in Section 2.1 of Reference 9 and summarized in Section 6.2. Applicability of the field equations to represent the processes and phenomena that can occur in the NPM is evaluated, along with an assessment of the ability of the NRELAP5 numerical solution to approximate the set of governing field equations. This evaluation addresses the mathematical models implemented in NRELAP5 for the NuScale LOCA analysis, and considers the applicability of the assumptions and processes involved in developing the NRELAP5 system of governing equations, and closure relations.

The numeric solution evaluation considers convergence, conservation of physical properties, and stability of code calculations performed to solve the set of governing equations for an NRELAP5 NPM model. The objective of this evaluation is to summarize information regarding the domain of applicability of the numerical techniques and user options that may impact the accuracy, stability, and convergence of NRELAP5 calculations. User guidelines for model development and execution were developed based on “lessons learned” during the code reviews and assessments. The guidelines include requirements for assuring convergence of solutions, accounting for uncertainty in results and monitoring code function to ensure that the basic conservation equations are being solved correctly.

8.3.1.1 Conservation of Mass, Momentum and Energy

NRELAP5 LOCA applications do not use the three-dimensional modeling capability of RELAP5-3D©. The one-dimensional equations and numerics have been used in versions of the RELAP5 codes for many years so their pedigree has been well established by code assessments and applications. The semi-implicit solution technique used by NuScale has been in the RELAP5 codes as the primary solution technique for the governing conservation equations since the initial development of the code. The solution technique continues to be used in NRELAP5 as discussed in Section 2.1.3 of the NRELAP5 code manual (Reference 9).

The basic governing equations for mass, momentum, and energy conservation use area-averaging for vapor and liquid fields. Mass, momentum, and energy conservation equations are written for each field, resulting in what is referred to as a six-equation model. The governing equations are discussed in Sections 6.2.1 through 6.2.4. The basic governing equations in NRELAP5 are generally accepted as reasonable representations of the applicable physical laws that govern the steady-state and transient behavior of thermal-hydraulic systems.

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Energy transfer into and out of the phases from the boundaries is governed by correlations discussed in Sections 6.2.5. Heat conduction within structures is modeled by the one-dimensional heat conduction equation discussed in Section 6.3.

Models are also included for trips and control systems as discussed in Section 6.5. This feature is used to model the safety-related system actuations, control power, set boundary conditions, determine ranges (minimum and maximum values) of selected variables including the FOMs, and other functions within the NRELAP5 models.

NuScale performed acceptance testing and procurement requirements as part of the commercial grade dedication of RELAP5-3D© to serve as the development platform for NRELAP5. The testing and inspection verified that RELAP5-3D© has the necessary critical characteristics to be used as the code development platform for NRELAP5. The critical characteristics include the suitability of the basic governing equations described above for the NuScale application.

8.3.1.2 Numerical Solution Techniques

The entire fluid domain of interest is divided into control volumes connected via junctions where the flow velocities are defined. The heat transfer into or out of control volumes are defined through heat structures where the heat conduction equation is solved considering the relevant heat transfer regime in the communicated control volume.

The difference equations implement mass and energy conservation by equating accumulation to the rate of mass or energy in through the cell boundaries, minus the rate of mass or energy out through the cell boundaries, plus source and sink terms. This approach necessitates defining mass and energy volume average properties and requiring knowledge of velocities at the volume boundaries. The velocities at the cell edges are defined through the use of momentum control volumes centered on the mass and energy cell boundaries. This approach results in a numerical scheme having a staggered spatial mesh with the momentum control volumes extending from the mass and energy cell centers to the neighboring mass and energy cell centers. The scalar properties of the flow (pressure, specific internal energies, and void fraction) are defined at mass and energy cell centers, while the vector quantities (velocities) are defined on the mass and energy cell boundaries.

The governing equations for the system model are solved numerically using a semi- implicit finite-difference technique. A nearly-implicit finite-difference technique which allows violation of the material Courant limit, is also available. However, the LOCA EM and the supporting assessment calculations use only the semi-implicit numerical scheme. The semi-implicit numerical solution scheme is based on replacing the system of differential equations with a system of finite difference equations partially implicit in time.

When generating a solution of finite difference equations, there is a possibility that the solution may not be converged. This could be the result of an ill-posed problem, inappropriate time step size selection, inadequate spatial nodalization, or an instability. Sensitivity studies have proven useful to ensure convergence and stability of the NRELAP5 solutions.

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Adherence to the modeling requirements of RELAP5 assist in ensuring that the governing equations are well posed. Requirements for nodalization and time step sensitivity studies comply with 10 CFR 50 Appendix K requirements and ensure converged solutions. Solutions are examined to identify unstable or unphysical behavior.

8.3.2 NuScale Facility Scaling

The NIST-1 facility is designed to simulate the integral system behavior of a single NPM immersed in a single bay within the reactor pool. The scaling analysis was performed to determine the geometric dimensions and operating conditions for the NIST-1 facility. The purpose of the scaling analysis was to design an IET facility that can be used to obtain quality data for thermal-hydraulic system safety analysis code validation. The hierarchical two-tiered scaling (H2TS) (Reference 99) method was used to perform the RCS natural circulation scaling and the scaling of LOCA and ECCS. The scaling analysis generated the sets of dimensionless groups that needed to be preserved to accurately simulate the high-ranked phenomena identified in the NuScale LOCA PIRT. The figures of merit were the peak CNV pressure and the collapsed liquid level above the top of the core. The scaling analysis also documented the scaling distortions between the NIST-1 facility and the NPM design, and evaluated the effects of these distortions.

Detailed documentation of the NIST-1 scaling analysis is available in the NIST-1 Facility Scaling Reports. Section 8.3.2.1 summarizes the scaling objectives and methodology. The approaches for RCS scaling natural circulation scaling and the scaling of LOCA and ECCS are briefly presented in Sections 8.3.2.2 and 8.3.2.3, respectively.

8.3.2.1 Scaling Objectives and Methodology

The general objective of the scaling analysis was to obtain the physical dimensions and operating conditions of a reduced-scale test facility capable of simulating the important flow and heat transfer behavior of a NPM under the LOCA conditions. To develop a properly scaled test facility, the following specific objectives were met for each operational mode of interest.

• The thermal-hydraulic processes that should be modeled were identified. • The similarity criteria that should be preserved between the test facility and the full- scale prototype were obtained. • The priorities for preserving the similarity criteria were established. • Specifications for the test facility design were established. • Biases due to scaling distortions were quantified. • The critical attributes of the test facility that must be preserved to meet testing requirements were identified.

Different similarity criteria were obtained for the different modes of system operation. These criteria depend on the geometry of the components, the scaling level required to

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address the transport phenomena of interest, and the initial and boundary conditions for each particular mode of operation.

To ensure that the scaling objectives were met in an organized and clearly traceable manner, a general design framework (GDF) was established. The model for this framework includes features drawn from the NRC severe accident scaling methodology presented in NUREG/CR-5809 (Reference 99). A flow diagram for the GDF is presented in Figure 8-7.

Figure 8-7. General design framework for the NuScale Integral System Test facility

Experimental Objectives

The first task outlined by the GDF was to specify the experimental objectives. The experimental objectives define the types of tests that will be performed to address specific design or certification needs. These objectives determined the general modes of operation that should be simulated in the test facility.

The objective of the NuScale LOCA test program was to obtain qualified data to benchmark the computer codes and models that will be used to evaluate the safety of the NPM. This includes: 1) measurements of transient and steady-state, single-phase

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natural circulation flow in the integrated RPV, and 2) characterization of the thermal- hydraulic phenomena in the RPV, containment, and CPV during the three periods of the LOCA.

Loss-of-Coolant Accident Phenomena Identification and Ranking Table

The second task outlined by the GDF was the development of a PIRT. The PIRT presented in Section 4 was used as the basis. The nature of scaling forbids exact similitude of all of the parameters of a reduced-scale test facility with those of a full-scale prototype. As a result, the design and operation of the test facility was based on simulating the thermal-hydraulic processes most important to the system operational modes that will be explored. The PIRT identified the different phases of a LOCA and most important thermal-hydraulic phenomena within those phases that should be simulated in the test facility. All of the highly ranked integral system phenomena identified in the NuScale LOCA PIRT are observed in the NIST facility to some degree. Although majority of the high-ranked PIRT phenomena are fully covered in NIST-1, the NIST facility is not the primary source of validation data for some phenomena. For example, the NIST facility does not model the details of the core fuel rods or core sub- channels; therefore, CHF data has been obtained in a separate full-scale test facility (see Section 7.3). Similarly, detailed information regarding the helical coil SG thermal- hydraulic performance is obtained from SIET TF-1 and TF-2 experiments (see Section 7.4).

Description of the H2TS Method

The third step in the GDF was to perform a scaling analysis for each of the hierarchical levels (e.g., systems and subsystems) and their modes of operation defined in the previous section. The H2TS method has been successfully used to develop the similarity criteria necessary to scale the APEX-600 and APEX-1000 systems for LOCA transients. The H2TS method was developed by the NRC and is fully described in Appendix D of NUREG/CR 5809 (Reference 99).

Figure 8-8 presents the four basic elements of the H2TS analysis method. The first element consists of subdividing the plant into a hierarchy of systems. Each system is subdivided into interacting subsystems which are further subdivided into interacting modules which are further subdivided into interacting constituents (materials) which are further subdivided into interacting phases (liquid, vapor or solid). Each phase can be characterized by one or more geometrical configurations and each geometrical configuration can be described by three field equations (mass, energy and momentum conservation equations). Each field equation can incorporate several processes. Figure 8-9 presents the breakdown of the NuScale system into hierarchical levels and high level processes to be scaled. It represents a roadmap used to structure the scaling analyses. The RCS and the ECCS were the focus of the scaling study.

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SYSTEM SCALE TOP-DOWN/SYSTEM SCALING BREAKDOWN IDENTIFICATION ANALYSIS

PROVIDE: PROVIDE HIERARCHY PROVIDE: FOR:  SYSTEM  CONSERVATION HIERARCHY  VOLUMETRIC EQUATIONS CONCENTRATIONS IDENTIFY:  AREA DERIVE: CONCENTRATIONS  SCALING GROUPS AND CHARACTERISTIC  PROCESS TIME CHARACTERISTIC TIME  CONCENTRATIONS SCALES RATIOS  GEOMETRIES  PROCESSES ESTABLISH:

 SCALING HIERARCHY

IDENTIFY:

 IMPORTANT PROCESSES FOR BOTTOM-UP SCALING ANALYSIS

BOTTOM-UP/PROCESS SCALING ANALYSIS

PERFORM:

 CONSERVATION EQUATIONSDETAILED SCALING ANALYSIS FOR IMPORTAT LOCAL PROCESSES

DERIVE AND VALIDATE:  SCALING GROUPS

Figure 8-8. Flow diagram for the hierarchical, two-tiered scaling analysis (NUREG/CR-5809)

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{{

}}2(a),(c) Figure 8-9. NuScale system breakdown into hierarchical levels and primary operational modes to be scaled

After identifying and subdividing the system of interest, the next step was to identify the scaling level at which the similarity criteria should be developed. This was determined by examining the phenomena being considered. For example, if the phenomenon being considered involves mass, momentum or energy transport between materials such as water and solid particles, then the scaling analysis would be performed at the constituent level. If the phenomenon of interest involves mass, momentum, or energy transport between vapor and liquid, then the scaling analysis would be performed at the phase level. Therefore, identifying the scaling level depends on the phenomenon being addressed.

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Thermal-hydraulic phenomena involving integral RCS interactions, such as primary system depressurization or loop natural circulation, would be examined at the “system” level. Thermal-hydraulic phenomena, such as SG heat transfer, would be examined at the “subsystem” level. Specific interactions between the steam-liquid mixture and the stainless steel structure would be examined at the “constituent” level.

The H2TS method required performing a “top-down” (system) scaling analysis. The top- down scaling analysis examines the synergistic effects on the system caused by complex interactions between the constituents deemed important by the PIRT. Its purpose is to use the conservation equations at a given scaling level to obtain characteristic time ratios and similarity criteria. It also identifies the important processes to be addressed in the bottom-up scaling analysis.

The H2TS method also required performing a “bottom-up” (process) scaling analysis. This analysis provides similarity criteria for specific processes such as flow pattern transitions and flow dependent heat transfer. The focus of the bottom-up scaling analysis is to develop similarity criteria to scale individual processes of importance to system behavior as identified by the PIRT.

Test Facility Specifications and Scaling Ratios

The fourth step of the GDF was to document all of the test facility design and operation specifications. All of the essential geometric features and operating parameters that must be carefully measured and documented to ensure accurate code simulations of the important thermal-hydraulic phenomena were identified and designated as critical attributes.

The NIST-1 facility was developed by modifying the existing MASLWR test facility at Oregon State University. This was accomplished by establishing a fixed set of scale factors for component lengths, flow areas, and volumes. These scale factors were obtained through an iterative process that included a practical assessment of component costs, ease of operation, material availability, and instrumentation accuracy for the scale selected. Having fixed the length, volume, and flow area scale factors for the test facility, and assuming fluid property similitude, the scaling ratios obtained using the governing equations for loop natural circulation were used to define the scale factors for the adjustable parameters. That is, the core power, component heat transfer areas, SG heat removal rate, and loop resistance were adjusted to preserve the requirement of isochronicity. Table 8-12 lists the required temporal and geometric scale factors for the NIST-1 facility under the requirement of isochronicity and fluid property similitude.

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Table 8-12. Scaling factors for NIST-1 facility

{{

}}2(a),(c)

The NIST-1 tests start from steady-state natural circulation conditions. {{

}}2(a),(c). The test facility operates near prototypic pressures and temperatures and operates with the same working fluid: water. Therefore, fluid property similitude is invoked. This means that the fluid property ratios are near to unity in all of the scale ratios, thereby simplifying the analysis.

8.3.2.2 Reactor Coolant System Natural Circulation Scaling

Figure 8-10 provides a flow diagram that describes the scaling analysis process for the RCS natural circulation operational mode. First, a top-down scaling analysis was performed. This included an analysis at the system level (integrated loop behavior) for normal operating conditions. {{

}}2(a),(c) Further details are available in the NIST-1 facility scaling reports.

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2(a),(c) }}

Figure 8-10. Scaling analysis flow diagram for single-phase primary loop natural circulation

8.3.2.3 Loss-of-Coolant Accident and Emergency Core Cooling System Scaling

The scaling analysis approach for LOCA and actuation of the ECCS includes the following four related scaling analyses:

• RCS depressurization • containment vessel pressurization • long-term recirculation cooling • Reactor Building pool heat-up

During ECCS operation, the RPV transports energy and mass to the containment. All of the mass and energy leaving the RPV is captured by the CNV. The CNV transports energy to the reactor pool.

The objectives of the LOCA/ECCS scaling analyses were to scale the

• RCS depressurization and containment pressurization behavior during the blowdown and venting phases of the LOCA.

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• RCS, containment cooling, and the reactor pool heat-up during the long-term recirculation cooling phase of the LOCA.

Reactor Coolant System Depressurization Scaling

Figure 8-11 shows top-down and bottom-up scaling analyses performed for RCS depressurization scaling. This included an analysis at the system level (integrated loop behavior) for RCS initial conditions at 50 percent power. {{

}}2(a),(c) {{

}}2(a),(c)

Figure 8-11. Scaling analysis flow diagram for reactor coolant system depressurization

Containment Pressurization Scaling

Top-down and bottom-up scaling analyses performed for the scaling of containment pressurization are shown in Figure 8-12. {{ . }}2(a),(c)

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{{

}}2(a),(c) {{

}}2(a),(c) Figure 8-12. Scaling analysis flow diagram for containment pressurization

Long-Term-Cooling Phase Scaling

Long-term recirculation occurs after the CNV and RPV pressures have become nearly equalized and the flow through the RRVs is from containment to RPV. As shown in Figure 8-13, top-down and bottom-up scaling analyses were performed for long-term phase scaling. {{ }}2(a),(c)

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{{

}}2(a),(c)

{{

}}2(a),(c) Figure 8-13. Scaling analysis flow diagram for long-term recirculation cooling mode

Reactor Pool Heatup Scaling

Heat transferred from the CNV exterior surface creates a heated plume of fluid that rises to the top of the pool. The heated plume mixes with the liquid at the top of the pool to create a thermally stratified layer. The thermally stratified layer consists of two regions; a well-mixed layer with uniform temperature at the surface of the pool and a partially mixed thermocline that extends downward and serves as a transition layer to the colder liquid in the pool. The thermal stratification layer grows over time. During the heat up of the pool, heat and mass are lost from the pool to the air at the pool interface due to evaporation. Heat is also transferred to the Reactor Building pool steel liner and concrete structures.

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Top-down and bottom-up scaling analyses performed for reactor pool heatup scaling are shown in Figure 8-14. {{

}}2(a),(c) {{

}}2(a),(c)

Figure 8-14. Scaling analysis flow diagram for Reactor Building pool heat-up

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8.3.2.4 As-Built NuScale Facility Scaling Summary

Comparison of plant behavior and the NIST test facility behavior for various events is presented in this section and in Section 8.3.4. Distortions exist between the plant and NIST IET as it does for any other scaled test facility. The purpose of the IET is to provide a scaled facility simulating the phenomena important to plant behavior with relative magnitudes that are similar to the plant for use in validating the models and integral behavior of the analysis code. The distortion analyses presented in this section for as- built NIST facility and in Section 8.3.4 for as-performed NIST tests show that the NIST facility meets these criteria. The top-down portion of the NIST-1 scaling analysis presented in Section 8.3.2.3 was expanded to perform an additional quantitative evaluation of the distortions in the as-built NIST-1 facility. The mass/energy balance equations were re-defined to include additional terms that better quantify the distortion in various phenomena seen in the RCS and CNV during a typical LOCA. The control volume balance equations derived for the RCS and CNV include

• {{

}}2(a),(c)

For quantifying the distortions, the following terms in the energy balance equations were explicitly accounted for in the top-down scaling analysis

• {{

}}2(a),(c)

The dimensionless forms of the mass/energy balance equations were derived by identifying the characteristic scales appearing in the balance equations. π groups characterizing the ratio of characteristic times for each process were defined based on the dimensionless equations.

Table 8-13, Table 8-14, and Table 8-15 summarize the mass flow paths and heat flow paths for the RCS and CNV considered in the top-down scaling analysis. The heat and mass flows into the control volume have a positive sign; whereas the negative sign represents heat and mass flow out of the control volume. Three mass flow rates are identified for both NPM and NIST-1 and are symmetric between the RCS and CNV. The same number of heat flow paths are identified for the RCS in NPM and NIST-1. {{

}}2(a),(c) Two major heat transfer paths are identified for the CNV: the first path is the heat transfer on the

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inner surface of the containment wall, the second is the heat transfer from the outer surface of the reactor pressure vessel. {{

}}2(a),(c)

A summary of π groups in the dimensionless mass/energy balance equations is given in Table 8-16 and Table 8-17. As depicted in Table 8-16 and Table 8-17, {{

}}2(a),(c)

Table 8-13. Mass Flow Paths for NPM and NIST-1 (RCS and CNV)

Flow Path No Description {{

}}2(a),(c)

Table 8-14. Heat Flow Paths for RCS in NPM and NIST-1

Heat Flow Path No Description {{

}}2(a),(c)

Table 8-15. Heat Flow Paths for Containment in NPM and NIST-1

Heat Flow Path No Description {{

}}2(a),(c)

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Table 8-16. Description of π Groups for the RCS Mass/Energy Balance

𝚷 Group Description {{

}}2(a),(c)

Table 8-17. Description of π Groups for the Containment Mass/Energy Balance

𝚷 Group Description {{

}}2(a),(c)

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𝚷 Group Description {{

}}2(a),(c)

These π groups were evaluated based on NRELAP5 simulations of the NPM and as-built NIST-1 facility for the following events:

• 100 percent discharge line break on the CVCS line (similar to NIST-1 HP-06 test)

• 100 percent high point vent line break (Similar to NIST-1 HP-07 test)

• Inadvertent opening of a single RVV (Similar to NIST-1 HP-09 test)

{{

}}2(a),(c) The key conclusions of this analysis are summarized below.

1. {{

}}2(a),(c)

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{{

}}2(a),(c)

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8. {{

}}2(a),(c) The scaling and distortion analysis methodology presented above is used to analyze the impact of biases in initial and boundary conditions and differences in operating procedures of the final NIST-1 IET data used in Section 8.3.4.

8.3.3 Assessment of NuScale Facility Integral Effects Test Data

The NIST-1 IET data that supports the validation of NRELAP5 for NPM LOCA analysis includes the following tests.

• {{

}}2(a),(c)

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{{

}}2(a),(c)

In addition to the above IETs, two more IETs were performed in the NIST-1 facility; HP- 43 (Inadvertent opening of RVV without DHRS) and HP-49 (Inadvertent opening of RRV without DHRS) (see Table 7-6). The results of NRELAP5 assessment against these test data are available in Appendices B and C. As shown in Sections 7.5.5 to 7.5.8 and Appendices A and B, in general, NRELAP5 predicted the NIST-1 IET data with excellent agreement. This shows that NRELAP5 is capable of predicting the phenomena and process occurring in the NIST-1 facility including system interactions. Further, evaluations of these assessments for each high-ranked PIRT phenomenon are summarized in Table 8-19.

8.3.4 Evaluation of NuScale Integral Effects Tests Distortions and NRELAP5 Scalability

The scaling and distortion analysis summarized in Section 8.3.2.4 identified and quantified scaling distortions in the as-built NIST-1 facility {{

}}2(a),(c)

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{{

}}2(a),(c)

The NuScale NRELAP5 LOCA EM was updated to simulate NIST-1 IETs HP-05, HP-06, HP-06b, HP-07, and HP-09 in the NPM. {{

}}2(a),(c)

The results showed that the biases, differences, and distortions between the NPM design and the NIST-1 facility can be accounted for using NRELAP5, and NRELAP5 is scalable to model phenomena and process in the NPM during LOCA events.

8.3.4.1 NuScale Facility Powered Natural Circulation Test (HP-05)

The NPM relies on natural circulation flow as the primary mechanism to remove energy produced in the core and to deposit that energy in the SG tubes. The core power provides the driving force, with resistance to the flow caused by form and friction losses along the primary coolant path. The NIST-1 test facility is a scaled model of the NPM that uses the same natural circulation mechanism to move energy from the core heater rods to the model SG tubes. Scaling factors for the NIST-1 facility are as shown in Table 8-12.

Test NIST-1 HP-05 was conducted to characterize the natural circulation flow rate and pressure drop in the NIST-1 test facility at various core power levels. As shown in Section 7.5.5 NRELAP5 predicted the HP-05 test data for the primary loop flow rate, core inlet temperature, and upper riser inlet temperature with reasonable-to-excellent agreement (see Figure 7-88 to Figure 7-90). These results demonstrate the applicability of NRELAP5 to predict the natural circulation flow in the NIST-1 facility over a range of power levels.

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{{

}}2(a),(c)

Figure 8-15 shows that the scaled NPM feedwater flow compares well with the test data. {{

}}2(a),(c) Figure 8-16 shows comparison of scaled NPM natural circulation flow to the test data. {{

}}2(a),(c)

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{{

}}2(a),(c),ECI

Figure 8-15. Comparison of HP-05 feedwater flow to test data {{

}}2(a),(c),ECI

Figure 8-16. Comparison of HP-05 reactor pressure vessel flow to test data

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The upper riser inlet temperature comparison to test data is shown in Figure 8-17. {{

}}2(a),(c)

In summary, the comparisons of scaled NRELAP5 calculations of the NPM RPV flow and fluid temperatures to NIST-1 NRELAP5 calculations and the test data indicate that the NIST facility is well scaled. {{

}}2(a),(c)

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{{

}}2(a),(c),ECI

Figure 8-17. Comparison of HP-05 upper riser inlet temperature to test data

{{

}}2(a),(c),ECI

Figure 8-18. Comparison of HP-05 core inlet temperature to test data

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8.3.4.2 NuScale Facility Loss-of-Coolant Accident and Inadvertent Reactor Vent Valve Opening Integral Effects Tests (HP-06, HP-07, and HP-09)

This section summarizes the results of the distortion analysis performed for the NIST-1 LOCA and inadvertent RVV opening IETs. The following initial/boundary condition biases, differences and scaling distortions between NPM and the as performed NIST-1 IETs have been identified to have a noticeable impact on the important LOCA parameters (i.e., RPV/CNV pressures and levels):

Initial Conditions:

The NIST-1 tests start from the steady-state natural circulation conditions. {{

}}2(a),(c)

The impact of bias in some of the initial conditions is summarized below:

Initial core power:

{{

}}2(a),(c)

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Initial RCS temperature/subcooling distribution:

{{

}}2(a),(c)

Initial reactor pool temperature:

{{

}}2(a),(c)

Reactor Core Power following Reactor Trip:

{{

}}2(a),(c)

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{{

}}2(a),(c)

CNV Wall Thickness and Material:

{{

}}2(a),(c)

Steam Generator Secondary Side Operation and Quantity of Steam Generators:

{{

}}2(a),(c)

NIST-1 CNV Shell:

{{

}}2(a),(c)

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{{

}}2(a),(c)

RPV Outside Surface Heat Transfer:

{{

}}2(a),(c)

RPV stored energy:

{{

}}2(a),(c)

8.3.5 Calculation of Peak CNV pressure

Since containment is an integral part of the NPM ECCS, Section 4.3 identifies peak containment pressure as one of the LOCA EM FOMs. However, as identified earlier in Sections 4.3, the peak containment pressure and temperature for containment performance are calculated with a different methodology (see the Containment Response Analysis Methodology - Reference 109). The top-down scaling analysis of Π groups representing the inventory and energy balance equations (see Section 8.3.2) can be used to provide more insights on the processes/phenomena governing the peak containment pressure. It was observed that the CNV pressurization during Phase 1a of the liquid space break is governed by {{

}}2(a),(c). As described in Section 9 of this report, the peak CNV pressure occurs following ECCS actuation in liquid space breaks. It was observed that the major processes that contribute to CNV pressurization during the early part of Phase 1b are {{ }}2(a),(c)

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{{

}}2(a),(c)

8.4 Summary of Adequacy Findings

8.4.1 Findings from Bottom-Up Evaluation

The bottom-up evaluation focused on determining the pedigree, applicability, fidelity to SET data, and scalability of the NRELAP5 closure relations and correlations that model the high-ranked phenomena as determined by the PIRT panel.

The pedigree of the identified closure relations and correlations was first established based on their historical development and subsequent assessment in the literature. Assessment cases were then identified to demonstrate the capability of NRELAP5 to predict the experimental data responses with reasonable-to-excellent agreement. Applicability of NRELAP5 to model the subject phenomena is established by demonstrating that the assessment cases cover the range of parameters that approximates the NPM range. The scalability evaluation was limited to whether the specific model or correlation is applicable for the NPM configuration over the range of conditions encountered in LOCA events.

Results of the bottom up evaluation are summarized in Table 8-18.

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Table 8-18. Summary of bottom-up evaluation of NRELAP5 models and correlations {{

}}2(a),(c)

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{{

}}2(a),(c)

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{{

}}2(a),(c)

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{{

}}2(a),(c)

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{{

}}2(a),(c)

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{{

}}2(a),(c)

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{{

}}2(a),(c)

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{{

}}2(a),(c)

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8.4.2 Findings from Top-Down Evaluation

Results of the adequacy evaluation based on the NIST-1 IETs are summarized in Table 8-19 below. All high-ranked phenomena are included in the table. Where the NIST-1 IETs do not provide information, or provide limited information, regarding NRELAP5 applicability to model the phenomenon an explanation is provided. Areas not covered, or partly covered, by the IETs are addressed by SETs or other means, e.g., sensitivity studies, bounding assumptions, component test data.

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Table 8-19. Applicability summary for high-ranked phenomena {{

}}2(a),(c)

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{{

}}2(a),(c)

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{{

}}2(a),(c)

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{{

}}2(a),(c)

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{{

}}2(a),(c) 8.4.3 Summary of Biases and Uncertainties

The NRELAP5 based LOCA EM was evaluated for applicability to analyzing LOCA events in the NPM. The applicability evaluation confirmed that the models and correlations in the NuScale LOCA EM are acceptable for simulating the important, i.e., high ranked, phenomena that determine the NPM response. Results of the LOCA EM applicability evaluation based on the bottom-up approach are summarized in Table 8-18. The overall evaluation of NRELAP5 applicability based on the top down approach is summarized in Table 8-19. The summaries in these tables show that the code is applicable for predicting LOCA response for the high-ranked phenomena that govern LOCA response in the NPM. A key element of the applicability confirmation is provided by SET and IET assessments that demonstrate reasonable-to-excellent agreement between NRELAP5 predictions and relevant experimental data.

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9.0 Loss-of-Coolant Accident Calculations

The primary purpose of the break spectrum calculations and sensitivity studies presented in this section is to support the development of the LOCA EM and to demonstrate its application for the evaluation of the NPM ECCS performance during postulated LOCAs. The specific objectives of this section are to:

• describe the progression of typical LOCA scenarios in the NPM with regard to the key phenomena and processes during different phases of the LOCA identified by the PIRT (Section 4.0), • present the results of the LOCA break spectrum calculations and other sensitivity calculations required by 10 CFR 50 Appendix K, and • present the results of additional sensitivity calculations that address the uncertainties in modeling of key phenomena affecting the LOCA progression. The initial/boundary conditions and inputs for key LOCA EM parameters used for this analysis are summarized in Appendix A.

9.1 Loss-of-Coolant Accident Progression in the NuScale Power Module

The LOCA progression for both a liquid and steam space break is presented in this section. A detailed discussion is provided for a 100 percent break of the RCS injection line and the high point vent line. As described in Section 4.2, the NPM LOCA has two distinct phases

1. A LOCA blowdown phase ( Phase 1a) begins with a postulated break in the RCS pressure boundary initiating a blowdown into the CNV and ends with opening the ECCS valves. 2. Phase 1b begins with the opening of ECCS valves resulting in pressure equalization between the RPV and CNV and the return of discharged fluid from the CNV to the RPV.

The long-term cooling phase begins when the pressure and level between the RPV and CNV stabilizes, and a stable natural recirculation flow pattern is established {{ }}2(a),(c)

The LOCA calculations are extended to {{ }}2(a),(c) following the flow reversal on the RRVs to ensure that the stable equilibrium collapsed levels are achieved in the riser. The LOCA scenarios described in the following sections assume full-break area, no loss of AC or DC power, no single failure, and do not credit either DHRS train. These conditions were chosen to represent a typical application of the conservative 10 CFR 50 Appendix K LOCA EM.

9.1.1 Liquid Space Break

The RCS injection line connects the CVCS system to the RPV riser section, and crosses the CNV (approximately {{ }}2(a),(c) above TAF inside the riser). A 100 percent break

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on the RCS injection line inside the CNV at time zero causes immediate choking at the break location as shown in Table 9-1. The mass and energy release into the CNV through the break results in rapid pressurization of the CNV and depressurization of the RPV. The MPS generates the reactor trip signal based on {{ }}2(a),(c). This signal is followed by CNV isolation with {{ }}2(a),(c) The reactor trip signal includes a {{ }}2(a),(c) delay to conservatively bound any {{ }}2(a),(c) The control rods drop to insert large negative reactivity and the drop is completed at approximately {{ }}2(a),(c) seconds. The containment isolation signal isolates {{ }}2(a),(c) If the DHRS was credited it would be activated at this time.

Phase 1a of the NPM LOCA includes the mass and energy release from the break location into the CNV and is terminated by the opening of the ECCS valves. For the 100 percent injection line break scenario, the ECCS valves are actuated on {{

}}2(a),(c). Figure 9-1 compares the break flow with the net ECCS valve flow during the transient. The RPV and CNV pressure responses shown in Figure 9-3 are a result of the behavior of each component of the energy balance shown in Figure 9-2. As shown, the energy release to the CNV through the break and ECCS valve flow is significantly larger than the energy release to the RPV by core heat transfer. Heat transfer from the CNV wall to the reactor pool causes a continuous depressurization of both RPV and CNV after the initial pressurization of the CNV. As shown in Figure 9-3, the peak containment pressure occurs at the time of the ECCS valve opening.

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Table 9-1. Event table for 100 percent reactor coolant system injection line break

{{

}}2(a),(c)

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{{

}}2(a),(c)

Figure 9-1. Break and emergency core cooling system valve flows to the containment vessel for 100 percent injection line break {{

}}2(a),(c) Figure 9-2. Integrated energy for 100 percent injection line break

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{{

}}2(a),(c)

Figure 9-3. Reactor pressure vessel and containment vessel pressure for 100 percent injection line break

As the RCS loses inventory, first through the injection line break and later through the ECCS valves, the collapsed level continuously drops, as shown in Figure 9-4 (see Section 5.1.2.6 for calculation of collapsed liquid level). After approximately {{ }}2(a),(c) seconds, the pressurizer is completely emptied (pressurizer level plotted on right side Y-axis of Figure 9-4) and the ECCS valves open when the collapsed liquid level is approximately {{ }}2(a),(c) above TAF. The collapsed liquid level drops another {{ }}2(a),(c) after the ECCS activation. Due to {{ }}2(a),(c), a portion of the RPV liquid inventory is temporarily {{ }}2(a),(c) at approximately {{ }}2(a),(c) seconds. As this inventory returns to the riser the RPV collapsed level increases. After the RPV and CNV pressure equilibrates, the pressurizer completely drains and an equilibrium fluid level is achieved at approximately {{ }}2(a),(c) above the TAF. The CCFL at pressurizer baffle plate is calculated to occur at 571 second and lasts for less than 1 second.

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{{

2(a),(c) }}

Figure 9-4. Comparison of collapsed liquid levels in reactor pressure vessel, containment vessel, and pressurizer for 100 percent injection line break

{{

}}2(a),(c)

The minimum core MCHFR is established shortly following the event initiation with slight reduction in magnitude with respect to its steady state value. {{

}}2(a),(c)

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{{ }}2(a),(c) Because the core remains covered and CHF is not observed, the peak cladding and fuel centerline temperatures remain cool (see Figure 9-7). The maximum cladding temperature {{ }}2(a),(c) and maximum fuel centerline temperature {{ }}2(a),(c) occur at {{ }}2(a),(c). Therefore, the 10 CFR 50.46 requirement of maximum allowed cladding temperature of 2200 degrees F is not challenged.

Phase 1b of the LOCA begins with the opening of the ECCS valves which produces the {{ }}2(a),(c) Stable natural circulation flow is established during this phase when the steam flowing into the CNV is condensed inside the CNV and condensate flow enters the RPV through the RRVs. The depressurization of both RPV and CNV continues as the heat transfer from the CNV wall to reactor pool is larger than the core decay power as shown in Figure 9-2. It is important to note that {{

}}2(a),(c) {{

}}2(a),(c) Figure 9-5. Core flow for 100 percent injection line break

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{{

}}2(a),(c)

Figure 9-6. Core minimum critical heat flux ratio during 100 percent injection line break

{{

}}2(a),(c) Figure 9-7. Peak cladding and fuel centerline temperature during 100 percent injection line break

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9.1.2 Steam Space Break

The largest steam space break occurs on the RCS high point vent line. The sequence of events is described in Table 9-2. Similar to the RCS injection line break discussed in the previous section, the 100 percent break on the high point vent line causes the immediate generation of the reactor trip signal based on the {{ }}2(a),(c) followed by a {{ }}2(a),(c) and a {{ }}2(a),(c) The control rods are fully inserted within approximately {{ }}2(a),(c) seconds of break initiation. Containment and secondary isolation occurs {{ }}2(a),(c) after the {{ }}2(a),(c) If DHRS were credited, it would be activated at this time.

{{

}}2(a),(c)

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Table 9-2. Event table for {{ }}2(a),(c) {{

}}2(a),(c)

On break initiation, the flow immediately chokes and remains choked for approximately {{ }}2(a),(c) seconds. The discharge of high-enthalpy steam from the RPV causes rapid depressurization of the RPV and pressurization of the CNV as shown in Figure 9-8. This immediately causes a reactor trip as shown in Table 9-2. {{

}}2(a),(c)

The collapsed level above the TAF decreases after the LOCA initiation as shown in Figure 9-10. {{

}}2(a),(c)

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{{

}}2(a),(c) The sensitivity results are discussed in Section 9.6.3.

As the RCS inventory is lost through the break, collapsed level continues to decrease. As condensate accumulates inside the CNV, {{

}}2(a),(c), which is similar to that of the injection line break discussed in Section 9.1.1.

The core MCHFR is not a concern as demonstrated in Figure 9-11. The minimum MCHFR is established at the transient initiation and is very close to the value corresponding to steady state. The MCHFR margin quickly increases with time due to power and flow mismatch. {{

}}2(a),(c) {{

}}2(a),(c) Figure 9-8. Comparison of pressure for injection line and high point vent line breaks

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{{

}}2(a),(c)

Figure 9-9. Break and emergency core cooling system valve flow during 100 percent high point vent line break {{

}}2(a),(c)

Figure 9-10. Collapsed liquid levels during 100 percent high point vent line break

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{{

}} 2(a),(c)

Figure 9-11. Minimum critical heat flux ratio during 100 percent high point vent line break

9.2 Break Size

In Section 9.1, the LOCA progression is described for two unique break locations (RCS injection line and high point vent line). The purpose of this section is to discuss the effect of break area on the LOCA FOMs. The spectrum of break areas for different break locations is summarized in Table 5-7. The justification of the selected matrix is discussed in Section 5.4.2. The maximum break area for the pressurizer spray supply line break is determined by the {{ }}2(a),(c) on the connecting piping. The minimum area for the liquid and steam space breaks is determined by examining a wide range of break areas such that limiting values for the NPM LOCA FOMs are obtained within the analyzed range. {{ }}2(a),(c)

The timing of events is directly affected by the break area through the choking flow rate at the break location. The break flow rate is proportional to the area for similar upstream conditions. The smaller break size results in slower depressurization and lower mass/energy loss. Therefore, transient times for smaller breaks are longer than the larger break sizes. For instance, events for the 10 percent break size take {{ }}2(a),(c) times as long as the maximum break size. The area ratios between break area and maximum break area described in Table 5-7 are used to scale the time in order to present the results of different break areas on the same time scale.

The CNV pressure (maximum of all CNV control volumes) as a function of scaled time with different break areas are presented for the RCS injection line, RCS discharge line, and high point vent line breaks in Figure 9-12, Figure 9-13, and Figure 9-14,

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respectively. Figure 9-15 shows peak CNV pressure as a function of break size for different break locations.

For the liquid space breaks (injection and discharge line breaks), the occurrence of peak CNV pressure coincides with the ECCS activation and the peak CNV pressures are very similar for break sizes down to {{ }}2(a),(c) of the full-size break (Figure 9-12, Figure 9-13, Figure 9-15). Furthermore, with discharge line break, the scaled time of the peak CNV pressure is very similar for break sizes down to {{ }}2(a),(c) percent of the full-size (Figure 9-13). The smaller liquid space break sizes are shown to produce smaller peak pressure values (Figure 9-15). As shown in Figure 9-12 and Figure 9-13, the magnitude of CNV pressure rise at the time of ECCS activation are very similar for all the liquid space break sizes. The peak CNV pressure values differ for smaller break sizes due to lower CNV pressurization rates as result of lower break energy release.

In contrast to the liquid space break, the high point vent line break with different break areas produces different peak CNV pressures (Figure 9-14 and Figure 9-15). {{

}}2(a),(c)

{{

}}2(a),(c)

Figure 9-12. Peak containment vessel pressure and collapsed level above top of active fuel for different reactor coolant system injection line break sizes

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{{

}}2(a),(c) Figure 9-13. Peak containment vessel pressure and collapsed level above top of active fuel for different reactor coolant system discharge line break sizes

{{

}}2(a),(c) Figure 9-14. Peak containment vessel pressure and collapsed level above top of active fuel for different high point vent line break sizes

Figure 9-12 to Figure 9-14 also show the collapsed liquid level above TAF in the RPV riser section as a function of scaled time for different break sizes and three break locations; RCS injection line, RCS discharge line, and high point vent line breaks. Figure 9-15 shows the minimum collapsed liquid level above TAF in the RPV riser section for different break locations as function of break size.

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The final equilibrium levels established at the end of the transient are independent of break size and locations. The equilibrium collapsed liquid level is directly related to the geometry of the RPV and CNV as well as the value of the equilibrium pressure between two vessels. For the high point vent line break cases, the minimum level is always the equilibrium level {{

}}2(a),(c)

The core MCHFR as a function break size is plotted in Figure 9-16 for both RCS injection and high point vent line breaks. As discussed earlier in Section 9.1 for the full- size liquid and steam space breaks, the CHFR margin rapidly increases following the event initiation and {{

}}2(a),(c) However, the potential for fuel heat-up is not of concern as no CHFR limit violation is observed in any of the break spectrum cases. {{

}}2(a),(c) Figure 9-15. Peak containment vessel pressure and minimum collapsed liquid level as a function of break location and size

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{{

2(a),(c) }} Figure 9-16. Minimum critical heat flux ratio for injection line (left) and high point vent line (right) breaks

9.3 Decay Heat Removal System Availability

In the previous sections, no credit is taken for the DHRS operation. The DHRS adds an additional heat sink capacity during the NPM LOCA that impacts primarily the smaller break sizes as shown in Figure 9-17 and Figure 9-18 for RCS injection line breaks. {{

}}2(a),(c) However, when the DHRS operation is taken into account, all break sizes behave similarly and minimum collapsed liquid levels are the same as the final equilibrium level for most all of the break sizes. Similar to the impact on the minimum collapsed levels, the MCHFR is defined by the hot assembly steady state value. As discussed previously, the NuScale LOCA EM does not take credit for the DHRS operation to introduce additional and significant conservatism in satisfying the LOCA FOMs. Consideration is given to the DHRS operation here only to confirm that more adverse conditions are not created when crediting the DHRS.

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{{

}}2(a),(c)

Figure 9-17. Reactor coolant system and containment pressures for reactor coolant system injection line break without decay heat removal system (left) and with decay heat removal system (right) {{

2(a),(c) }}

Figure 9-18. Collapsed liquid level for reactor coolant system injection line break without decay heat removal system (left) and with decay heat removal system (right)

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9.4 Power Availability

The discussion presented previously assumes that both AC and DC power are available during the NPM LOCA. Loss of power is considered by assuming either loss of only AC power or loss of both AC and DC power. Figure 9-19 demonstrates that the loss of both AC and DC power has significant impact on peak containment pressure for the steam space breaks, but has minimal impact on liquid space breaks. The loss of all power causes an immediate reactor trip and de-energizes the ECCS valves. {{

}}2(a),(c)

{{

}}2(a),(c)

Figure 9-19. Effect of power availability on peak containment vessel pressure for injection line (left) and high point vent (right) line breaks

9.5 Single Failure

In all of the previous discussion, no single failure is assumed. As discussed in Section 5.4.3, the following single failures are considered in this section:

• failure of a single RVV to open, • failure of a single RRV to open, and • failure of one ECCS division (i.e., one RVV and one RRV) Figure 9-20 demonstrates that the single failures listed above have negligible impact on the peak containment pressure. In fact, the opening failure of a single RVV produces

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slightly smaller peak pressures inside the containment. The peak containment pressure of approximately {{ }}2(a),(c) psia is not affected by the single failure assumptions for the liquid space break. Therefore, the simulations with no single failure produce similar or conservative peak containment pressures for different break sizes. Similar conclusions can also be reached for the minimum collapsed levels in the RCS as demonstrated in Figure 9-21 where failure of a single RRV produces slightly better minimum collapsed levels above the TAF. In conclusion, a single failure based on failure to open an RVV and RRV does not produce more conservative results on peak CNV pressure and minimum collapsed level. Similar to the previous discussion, no CHF violation is calculated with the single failures considered as part of the NPM LOCA break spectrum. {{

}} 2(a),(c)

Figure 9-20. The effect of single failure on peak containment vessel pressure for reactor coolant system injection line (left) and high point vent (right) line breaks

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{{

}}2(a),(c) Figure 9-21. The effect of single failure on minimum collapsed level for reactor coolant system injection line (left) and high point vent line (right) breaks

9.6 Sensitivity Studies

Several sensitivity studies are performed to establish the basis for the NuScale LOCA EM. The sensitivity calculations are performed to address the effects of the modeling parameters such as nodalization, time-step size selection, CCFL behavior at the pressurizer baffle plate, ECCS valve parameters (such as IAB release pressure differential threshold, size/capacity, as well as valve stroke time). An additional sensitivity study is performed on core power distribution addressing the effects of core axial power shape and radial peaking assigned to the hot assembly. The sensitivity calculations are also performed to determine the impact of initial reactor cooling pool temperature. Justifications for other initial and boundary conditions selected for the conservative LOCA analysis are provided in Section 5.3.

9.6.1 Model Nodalization

Performing a nodalization sensitivity study is important to determine its impact on the key LOCA FOMs such as peak containment pressure and collapsed liquid level above TAF in the RPV riser. As described in Section 5.1, the NRELAP5 model uses one- dimensional components. In order to address the impact of nodalization on the NPM LOCA behavior, three nodalization schemes that conform to general NRELAP5 modeling guidelines are selected as shown in Table 9-3.

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Table 9-3. Number of volumes in reactor pressure vessel and containment vessel nodalization {{

}}2(a),(c)

The full range of break sizes for both RCS injection line and high point vent line breaks with three nodalization schemes are investigated. Both break locations are considered without DHRS operation, no loss of power, and no single failure.

Figure 9-22 shows the RPV and CNV pressures and collapsed liquid level above TAF in RPV riser for the RCS injection break with 100 percent break area without DHRS, no loss of power, and no single failure. The same parameters are plotted in Figure 9-23 for the RCS injection break with 10 percent break area without DHRS, no loss of power, and no single failure. With three different nodalization schemes, two key LOCA FOMs are shown to be similar including timing of event during the transient, {{

}}2(a),(c)

The results shown in Figure 9-24 for the 100 percent high point vent line break cases show the similarities. {{

}}2(a),(c) three different nodalization schemes provide similar LOCA response in RPV and CNV pressures and collapsed levels for the high point vent line break scenario.

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{{

}}2(a),(c)

Figure 9-22. Reactor pressure vessel and containment vessel pressure (left) and collapsed level above top of active fuel (right) for 100 percent reactor coolant system injection line break {{

}}2(a),(c) Figure 9-23. Reactor pressure vessel and containment vessel pressure (left) and collapsed level above top of active fuel (right) for 10 percent reactor coolant system injection line break

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{{

}}2(a),(c) Figure 9-24. Reactor pressure vessel and containment vessel pressure (left) and collapsed level above top of active fuel (right) for 100 percent high point vent line break

The hot assembly mass flux and core-wide MCHFR during RCS injection line break are shown in Figure 9-25 for the three nodalization schemes. The initial core MCHFR at the beginning of the transient is not affected by the number of hydrodynamic volumes in the NPM core, {{ }}2(a),(c) Furthermore, very similar MCHFRs are calculated for the {{ }}2(a),(c) correlation. When the hot channel assembly flow goes {{ }}2(a),(c) correlation that includes the {{ }}2(a),(c) is used. {{

}}2(a),(c) Therefore, the core nodalization has no material impact on predicting the CHF margin during a postulated NPM LOCA. However, a small difference is observed in the initial hot assembly flows with coarse and coarser nodalization. As described in Table 9-3, both coarse and coarser nodalization schemes use a coarse representation of the core and steam generators. As a result, the steady state natural circulation flow rate is slightly different when compared to the finer nodalization due to relatively small shift in natural circulation loop thermal center.

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{{

}}2(a),(c)

Figure 9-25. Hot channel core flow (left) and core critical heat flux ratio (right) during 100 percent reactor coolant system injection line break

9.6.2 Time-Step Size Selection

The NRELAP5 NuScale LOCA EM uses a semi-implicit numerical scheme with implicit coupling of the hydrodynamic and heat conduction solutions. The time-step size is restricted by the courant time-step size and the accumulation of the mass-error during the time integration. In general, the NPM LOCA simulations have a courant time-step size at approximately {{ }}2(a),(c) In order to address the effect of time-step size selection on the key NPM LOCA FOMs, various fractions of the problem courant time-step size are examined as shown in Figure 9-26 through Figure 9-29 for full size injection line and high point vent line breaks. For multipliers above approximately {{ }}2(a),(c) the max time-step size allowed for the calculations is mainly determined by the mass-error management. The figures show that the containment and RPV pressures, minimum collapsed level above the TAF in the RPV riser, hot channel mass flux, and hot channel MCHFR are all independent of the time- step sizes selected for the simulation.

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{{

2(a),(c) }}

Figure 9-26. Time-step size sensitivity on reactor and containment vessel pressures and reactor pressure vessel collapsed liquid level for 100 percent reactor coolant system injection line break. {{

}}2(a),(c) Figure 9-27. Time-step size sensitivity on hot assembly flow and minimum critical heat flux ratio for 100 percent reactor coolant system injection line break

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{{

2(a),(c) }} Figure 9-28. Time-step size sensitivity on reactor and containment vessel pressures and reactor pressure vessel collapsed liquid level for 100 percent high point vent line break

{{

}} 2(a),(c) Figure 9-29. Time-step size sensitivity on hot assembly flow and minimum critical heat flux ratio for 100 percent high point vent break

9.6.3 Counter Current Flow Limitation Behavior on Pressurizer Baffle Plate

{{ }}2(a),(c) A few of the break spectrum cases activated the CCFL flag at the pressurizer baffle plate, which did not allow liquid to readily drain from the pressurizer to the downcomer in the presence of upward steam flow. These break cases were limited to

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the larger pressurizer spray and vent line breaks. A study was performed {{

}}2(a),(c)

{{

}} 2(a),(c)

Figure 9-30. Effect of counter current flow limitation line slope on levels for 100 percent high point vent line break

9.6.4 Emergency Core Cooling System Valve Parameters

Operation of the ECCS valves varies based on the valve characteristics. The NPM ECCS valve specification provides minimum and maximum valve sizes and a range of differential pressures at which the IAB arming valve closes (locks) and opens (releases). A study was performed with liquid and steam breaks to evaluate separate and combined effects of the range of these valve characteristics on the LOCA FOMs.

Figure 9-31 shows the effect of IAB release pressure on peak CNV pressure and minimum collapsed liquid level as function of break size for injection line break. Since the large break size results in relatively rapid RCS depressurization, {{ }}2(a),(c)

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{{

}}2(a),(c)

Figure 9-32 and Figure 9-33 show the effect of RRV and RVV sizes on peak CNV pressure and minimum collapsed liquid level as function of break size. Overall the impact on ECCS valve size on peak CNV pressure and collapsed liquid level is {{ }}2(a),(c). Figure 9-32 shows only {{ }}2(a),(c) in minimum collapsed liquid level with {{ }}2(a),(c).

Conclusions of this study show that the {{ }}2(a),(c)

{{

}}2(a),(c)

Figure 9-31. Effect of inadvertent actuation block release pressure on peak containment vessel pressure and minimum collapsed liquid level above top of active fuel as a function of break size for reactor coolant system injection line break

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{{

}}2(a),(c) Figure 9-32. Effect of reactor recirculation valve size on peak containment vessel pressure and minimum collapsed liquid level for reactor coolant system injection line break {{

}}2(a),(c) Figure 9-33. Effect of reactor vent valve size on peak containment vessel pressure and minimum collapsed liquid level for reactor coolant system injection line break

9.6.5 Initial Reactor Pool Temperature

The maximum initial reactor cooling pool temperature of {{ }}2(a),(c) is used in the LOCA break spectrum calculations as discussed in the previous sections. Sensitivity studies covering the range of initial pool temperatures are performed to

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investigate the impact on the NuScale LOCA EM FOMs. Reactor pool temperatures ranging from {{ }}2(a),(c) to {{ }}2(a),(c) are considered. The RCS injection line break with sizes down to {{ }}2(a),(c) of the full-break size break area are analyzed. The peak CNV pressure and the minimum collapsed liquid level above TAF as a function of break size are plotted for the pool temperatures of {{ }}2(a),(c) in Figure 9-34. The effect of the initial pool temperature on the peak CNV pressure is more pronounced at {{ }}2(a),(c). Figure 9-35 compares the various components of the RPV and CNV energy balance for 100 percent (left figure) and 10 percent (right figure) injection line breaks. {{

}}2(a),(c) For all the initial pool temperatures investigated in the sensitivity calculation, no CHF violation is observed; therefore, the minimum MCHFR is defined by a value close to the steady state value. {{

}}2(a),(c)

Figure 9-34. Effect of initial reactor pool temperature on peak containment vessel pressure and minimum collapsed liquid level above top of active fuel for reactor coolant system injection line break

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{{

}}2(a),(c) Figure 9-35. Containment vessel to pool energy transfer at different initial pool temperatures for 100 percent (left) and 10 percent (right) reactor coolant system injection line break

9.6.6 Core Power Distribution

The sensitivity study is performed based on a full-range of break sizes for the RCS injection line break for the core power distribution considering:

• Generic axial power shapes to bound the axial peakings • {{ }}2(a),(c) core channel

Generic axial power shapes as shown in Figure 9-36 are used to investigate the effect on the key LOCA behavior and FOMs. The axial power shapes are chosen to represents a typical {{ }}2(a),(c). The {{ }}2(a),(c) shown in Figure 9-36 is used for all the LOCA calculations performed in this report. The axial peaking is determined to bound the values observed in the NPM core design (Appendix A).

{{

}}2(a),(c)

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{{

}}2(a),(c) Figure 9-36. Generic normalized axial power shapes

The RCS injection line break with full break area spectrum is analyzed without DHRS operation, no power loss, and no single failure. Figure 9-37 compares the RPV and CNV pressures and collapsed liquid level above TAF for different axial power shapes for RCS injection line break. Figure 9-38 shows the impact of axial power shapes on peak CNV pressure and minimum collapsed liquid level as function of different injection line break sizes. {{

}}2(a),(c)

Figure 9-39 shows the impact of axial power shapes on the hot assembly mass flux and MCHFR {{

}}2(a),(c) However, in all of the cases analyzed, the minimum core MCHFR transient value is close to the value determined at the initiation of the event.

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{{

}}2(a),(c) Figure 9-37. Effect of axial power shape on reactor pressure vessel and containment pressures and collapsed liquid level above top of active fuel for reactor coolant system injection line break

{{

}}2(a),(c) Figure 9-38. Effect of axial power shape on peak containment vessel pressure and minimum collapsed liquid level above top of active fuel for reactor coolant system injection line break

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{{

}}2(a),(c) Figure 9-39. Effect of axial power shape on hot assembly flow and minimum critical heat flux ratio during reactor coolant system injection line break

9.7 Loss-of-Coolant Accident Calculation Summary

The following conclusions are reached based on the beak spectrum calculations and sensitivity studies:

1. The core MCHFR rapidly increases following the initiation of a LOCA due to power/flow mismatch – power decreases faster than flow due to differences in process time constants. 2. {{

}}2(a),(c) In conclusion, there is no fuel CHF and hence no fuel heat-up for a NPM LOCA. 3. The most sensitive LOCA analysis parameters were determined to be DHRS unavailability (conservatively assumed), {{ }}2(a),(c) 4. {{

}}2(a),(c) 5. For all the break cases and sizes, the {{ }}2(a),(c) collapsed RPV level above TAF converges to approximately {{ }}2(a),(c) This value is independent of the LOCA progression. This value is directly related to the {{ }}2(a),(c) and the initial mass/energy inventory.

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6. Minimum RPV collapsed level during NPM LOCA transient is invariant of the break size down to {{

}}2(a),(c) 7. {{

}}2(a),(c) 8. {{

}}2(a),(c) 9. {{

}}2(a),(c)

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10.0 Conclusions

The NPM is unique when compared to any current operating power plant. It is based on an integral PWR design without coolant loops, coolant pumps, or pressurizer surge lines. The primary reactor system is driven by natural circulation with few connecting pipes and a simple safety-related system to mitigate the consequences of postulated accidents. In particular, the NPM design is not significantly challenged by LOCA events as primary system coolant is captured completely by the CNV, cooled, and returned to the RPV using a large reactor pool as the ultimate heat sink, which can provide cooling for many weeks.

The LOCA EM uses a conservative bounding approach to analyzing LOCA transients. It adheres to the relevant requirements of 10 CFR 50 Appendix K and follows the EMDAP described in RG 1.203. Multiple layers of conservatism are incorporated in the LOCA EM to ensure that a conservative analysis result is obtained. These conservatisms stem from application of the relevant modeling requirements of 10 CFR 50 Appendix K and through a series of conservative modeling features that have been incorporated.

The methodology uses the proprietary NRELAP5 computer code. RELAP5-3D© was procured and commercial grade dedication was performed as part of the procurement process by NuScale to establish the baseline NRELAP5 code for development. Subsequently, features were added and changes made to NRELAP5 to address the unique aspects of the NPM design and licensing methodology. The models and correlations used in the NRELAP5 code have been reviewed and, where appropriate, modified for use within the NuScale LOCA EM. Features added and changes made to address unique aspects of the NPM design and NuScale LOCA EM that applies to 10 CFR 50 Appendix K include the following:

• helical coil SG heat transfer and pressure drop models • core CHF models • wall condensation models • critical flow models • interfacial drag models for large diameter pipes

The NRELAP5 code includes all of the necessary models for characterization of the NPM hydrodynamics, heat transfer between structures and fluids, modeling of fuel, reactor kinetics models, and control systems. The geometry of certain NPM components dictates the use of specific correlations, {{ }}2(a),(c) The CHF correlations chosen to assess fuel conditions were selected based on full-scale fuel bundle performance tests over the range of conditions (flows, temperatures, and pressures) anticipated in the NPM during a LOCA event.

A number of conservatisms are built into the NuScale LOCA EM to ensure that conservative analysis results are obtained. Not only are applicable 10 CFR 50 Appendix

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K conservatisms present, but additional conservatisms above and beyond 10 CFR 50 Appendix K have been incorporated.

Conservatisms that are in addition to the 10 CFR 50 Appendix K requirements include:

• Not crediting actuation of DHRS in the break spectrum assessment to conservatively reduce the heat removed from the RPV • {{

}}2(a),(c)

A PIRT was developed that identified all of the important phenomena that could occur during a LOCA event. Phenomena and process ranking was performed in relation to specified FOMs, as described by RG 1.203. The PIRT also established a knowledge ranking for each of the phenomena identified. Using these FOMs, 21 phenomena were identified as important to correctly capture in the LOCA EM.

Extensive NRELAP5 code validation has been performed to ensure that the LOCA EM is applicable for all important phenomena and processes over the range of conditions encountered in the NPM LOCA. The validation suite includes many legacy SETs and IETs, as well as many SETs and IETs developed and run specifically for the NPM application. The SETs run for the NPM application were performed at the SIET facility on a model helical coil SG and at the Stern facility to obtain CHF data on a full-scale rod bundle test section. Integral effects tests were performed at the NIST-1 facility, a scaled representation of the complete NPM primary and secondary systems, as well as the CPV.

The EMDAP requires an applicability demonstration of the NRELAP5 code and tests. A unique aspect of the EMDAP applicability demonstration is the comparison of NRELAP5 simulations of LOCA events to NIST-1 test data and NRELAP5 simulation of the same LOCA event in the NPM. {{ }}2(a),(c) The reasonable-to-excellent agreement obtained by these comparisons establishes both the fidelity of the NIST-1 design to the NPM, and the applicability of NRELAP5 to accurately predict LOCA phenomena at both the NIST-1 and NPM scales. Limitations and modeling

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requirements were determined in this assessment process and are accounted for in the application of the LOCA methodology.

This topical report provides an example application of the LOCA EM in order to aid the reader’s understanding of the context of the application of the NuScale LOCA EM. These calculations are presented for break spectra that cover a range of break locations, break sizes, single failures, equipment unavailability, and initial and boundary conditions. The nodalization and time-step sensitivity required by 10 CFR 50 Appendix K and additional sensitivity calculations that address the uncertainties in modeling of key phenomena are performed. The analyses conducted demonstrate that the NPM retains sufficient water inventory in the primary system such that the core does not uncover or experience a CHF condition during a LOCA such that the minimum CHF ratio is greater than the analysis limit of {{ }}2(a),(c) as described in Section 7.3.6, and that containment design pressure is not challenged. The PCT is shown to occur at the beginning of the LOCA event and cladding temperature decreases as the transient evolves. Because no fuel heatup occurs for any design-basis LOCAs, the following regulatory acceptance criteria from 10 CFR 50.46 are met:

• PCT remains below 2,200 degrees Fahrenheit (1,204 degrees Celsius). • Maximum fuel oxidation is less than 0.17 times total cladding thickness. • Maximum hydrogen generation is less than 0.01 times that generated if all cladding were to react. • Coolable geometry is retained.

The methodology in this report is also used to support other analyses including:

1) events as described in Topical Report TR-0516-49416-P, “Non-Loss of Coolant Accident Methodology,”

2) containment peak pressure analysis as described in Technical Report TR-0516- 49084-P, “Containment Response Analysis Methodology,”

3) long term cooling as described in Technical Report, TR-0919-51299-P, “Long-Term Cooling Methodology,” and

4) inadvertent Opening of Reactor Pressure Vessel (RPV) Valves, including ECCS valves as described in Appendix B of this report, “Evaluation Model for Inadvertent Opening of RPV Valves”.

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11.0 References

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37. Colburn, A.P., and O.A. Hougen, "Design of Cooler Condensers for Mixtures of Vapors with Noncondensing Gases," Industrial and Engineering Chemistry, (1934): 26:1178- 1182. 38. Green, D.W. and Perry, R.H., Perry's Chemical Engineers' Handbook, 8th Edition, McGraw-Hill, New York, NY, 2008. 39. {{ }}2(a),(c) 40. Kataoka, I., Ishii, M., "Drift Flux Model for Large Diameter Pipe and New Correlation for Pool Void Fraction", International Journal of Heat and Mass Transfer, (1987): Vol. 30, No. 9, pp. 1927-1939. 41. Churchill, S.W., and H.H.S. Chu, "Correlating Equations for Laminar and Turbulent Free Convection From a Vertical Plate," International Journal of Heat and Mass Transfer, (1975): 18:1323-1329. 42. Rouhani, S.Z., "Modified Correlations for Void and Pressure Drop," AB Atomenergi, Sweden, Internal Report AE-RTC 841, March 1969. 43. Reserved. 44. Draft American Nuclear Society, “Decay Energy Release Rate Following Shutdown of Uranium-Fueled Thermal Reactors,” Proposed Standard ANS 5.1, LaGrange Park, IL, October 1973. 45. Ishii, M., "One-Dimensional Drift-Flux Model and Constitutive Equations for Relative Motion between Phases in Various Two-Phase Flow Regimes", Report No. ANL-77-47, Argonne National Laboratory, October 1977. 46. American Nuclear Society, “Decay Energy Release Rate Following Shutdown of Uranium-Fueled Thermal Reactors,” Proposed Standard ANS 5.1, LaGrange Park, IL, October 1971 (revised October 1973). 47. Shure, K., “Fission-Product Decay Energy,” WAPD-BT-24, Westinghouse Atomic Division, Bettis, December 1961. 48. American National Standards Institute/American Nuclear Society, “Decay Heat Power in Light Water Reactors,” ANSI/ANS-5.1-1979, LaGrange Park, IL. 49. American National Standards Institute/American Nuclear Society, “Decay Heat Power in Light Water Reactors,” ANSI/ANS-5.1-1994, LaGrange Park, IL. 50. American National Standards Institute/American Nuclear Society, “Decay Heat Power in Light Water Reactors,” ANSI/ANS-5.1-2005, LaGrange Park, IL. 51. Healzer, J.M., J.E. Hench, E. Janssen, and S. Levy, “Design Basis for Critical Heat Flux Condition in Boiling Water Reactors,” APED-5186, GE Company Private Report, July 1966. 52. General Electric, “General Electric BWR Thermal Analysis Basis (GETAB): Data, Correlation and Design Application,” NEDO-10958-A, 1977. 53. Biasi, L., et.al., “Studies on Burnout Part 3- A New Correlation for Round Ducts and Uniform Heating and Its Comparison with World Data,” Energy Nucl., (1967), 14:530-7. 54. Electric Power Research Institute, “Parametric Study of CHF Data,” Vol. 2, Palo Alto, California, 1983. 55. {{ }}2(a),(c)

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56. Zuber, N., “Hydrodynamic Aspects of Boiling Heat Transfer,” Ph.D. thesis, University of California at Los Angeles, 1959. 57. Groeneveld, D.C., et al., “Lookup Tables for Predicting CHF and Film-Boiling Heat Transfer: Past, Present and Future,” Nuclear Technology, (2005): 152:87-104. 58. Ferrell, J.K., and J. W. McGee, “Two-Phase Flow Through Abrupt Expansions and Contractions, Final Report, “Study of Convection Boiling Inside Channels”, TID-23394 (Vol. 3), June 1966. 59. U.S. Nuclear Regulatory Commission, "BWR Refill-Reflood Program Task 4.8-Model Qualification Task Plan," NUREG/CR-1899, August 1981. 60. Kim, S.J., “Turbulent film condensation of high pressure steam in a vertical tube of Passive Secondary Condensation System," PhD thesis, Korea Advanced Institute of Science and Technology, 2000. 61. Nylund O., et.al, “Hydrodynamic and heat transfer measurements on a full-scale simulated 36-rod BHWR fuel element with non-uniform axial and radial heat flux distribution”, 1970. 62. U.S. Nuclear Regulatory Commission, “Analysis of the FLECHT SEAST Unblocked Bundle Steam Cooling and Boiloff Tests,” NUREG/CR-1533, January 1981. 63. U.S. Nuclear Regulatory Commission, “PWR FLECHT SEASET Unblocked Bundle, Forced and Gravity Reflood Task Data Report,” NUREG/CR-1532, Volume 2, Appendix C, September 1981. 64. U.S. Nuclear Regulatory Commission, “Results of the Semiscale MOD-2A Natural Circulation Experiments,” NUREG/CR-2335, September 1982. 65. Reserved. 66. Electric Power Research Institute, “Final Report, Two-Phase Jet Modeling and Data Comparison,” EPRI NP-4362, March 1986. 67. U.S. Nuclear Regulatory Commission, "Countercurrent Flow of Air/Water and Steam/Water Flow above a Perforated Plate," NUREG/CR-1808, November 1980. 68. U.S. Nuclear Regulatory Commission, "Countercurrent Steam/Water Flow Above a Perforated Plate-Vertical Injection of Water," NUREG/CR-2323, September 1981. 69. Bankoff, S.G., R.S. Tankin, M.C. Yuen, and C.L. Hsieh, "Countercurrent Flow of Air Water and Steam/Water Through a Horizontal Perforated Plate", International Journal of Heat and Mass Transfer, (1981): Vol. 24, No. 8 pp 1381-1395. 70. U.S. Nuclear Regulatory Commission, “The Marviken Full Scale Critical Flow Tests, Summary Report, Joint Reactor Safety Experiments in the Marviken Power Station Sweden,” NUREG/CR-2671, May 1982. 71. The Marviken Full Scale Critical Flow Tests, Results from Test 22, Joint Reactor Safety Experiments in the Marviken Power Station Sweden, MXC-222, September 1979. 72. Modro, S.M., et. al., “Multi-Application Small Light Water Reactor Final Report,” INEEL/EXT-04-01626, Idaho National Engineering and Environmental Laboratory, December 2003. 73. Idelchik, I.E., “Handbook of Hydraulic Resistance,” Hemisphere Publishing, New York, NY, 3rd Edition. 74. Fletcher, C.D., et al., “Adequacy Evaluation of RELAP5/MOD3, Version 3.2.1.2 for Simulating AP600 Small-Break Loss-of-Coolant Accidents,” INEL-96/0400, April 1997.

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75. NuScale Power, LLC, “Topical Report: Subchannel Analysis Methodology,” TR-0915- 17564-P-A, Rev. 2, February 2019. 76. The RELAP5-3D Code Development Team, "RELAP5-3D Code Manual, Volume III: Developmental Assessment", INEEL-EXT-98-00834, Revision 4.1, October 2013. 77. Jeandey, C., et al., "Auto Vaporisation D'Ecoulements Eau/Vapeur, Departement des Reacteurs a Eau Service des Transferts Thermiques (Centre D'Etudes Nucleaires de Grenoble)," Report T.T. No. 163, July 1981. 78. U.S. Nuclear Regulatroy Commission, "Assessment of Two-Phase Critical Flow Models Performance in RELAP5 and TRACE against Marviken Critical Flow Tests," NUREG/IA- 0401, February 2012. 79. Elias, E., and G. S. Lellouche, "Two-Phase Critical Flow," International Journal of Multiphase Flow, (1994): Vol. 20, No. 91-168. 80. US Nuclear Regulatory Commission, “TRACE V5.0 Theory Manual – Volume 1: Field Equations, Solution Methods, and Physical Models,” June 2008, Agencywide Document Access and Management System (ADAMS) Accession No. ML120060218. 81. RELAP5 MOD3.3 Code Manual, Volume IV: Models and Correlations, October 2010. 82. {{

}}2(a),(c) 83. Lee, K.W., H.C. No, and C.H. Song, "Onset of Water Accumulation in the Upper Plenum with a Perforated Plate," Nuclear Engineering and Design, (2007): 237:1088-1095. 84. Wallis, G.B., One-dimensional Two-Phase Flow, McGraw-Hill, New York, NY, 1969. 85. Ilic, V., S. Banerjee, and S. Behling, ”Qualified Database for the Critical Flow of Water, Final Report”, EPRI-NP-4556, May 1986. 86. Zuber, N., and J. A. Findlay, “Average Volumetric Concentrations in Two-Phase Flow Systems”, Transactions of the ASME, Journal of Heat Transfer, (1965): 87:453-568. 87. The RELAP5-3D Code Development Team, "RELAP5-3D Code Manual, Volume III: Developmental Assessment", INEEL-EXT-98-00834, Revision 4.1, October 2013. 88. McAdams, W.H., Heat Transmission, 3rd Edition, McGraw-Hill, New York, NY, 1954. 89. Minkowycz, W.J., and E. M., Sparrow, “Local Nonsimilar Solutions for Natural Convection on a Vertical Cylinder,” Journal of Heat Transfer, (1974): 96(2), 178-183. 90. Plesset, M.S., and S. A. Zwick, "The Growth of Vapor Bubbles in Superheated Liquids", Journal of Applied Physics, (1954): 25, 493. 91. Lee, K., and D. J. Ryley, "The Evaporation of Water Droplets in Superheated Steam", Transactions of the ASME, Journal of Heat Transfer, (1968): pp. 445-451. 92. Aumiller, D.L., “The Effect of Nodalization on the Accuracy of the Finite-Difference Solution of the Transient Conduction Equation”, 2000 RELAP5 International Users Seminar, Jackson Hole, Wyoming, September 12-14, 2000. 93. Electric Power Research Institute, "The Chexal-Lellouche Void Fraction Correlation for Generalized Applications," NSAC-139, April 1991. 94. Inayatov, A.Y., "Correlation of Data on Heat Transfer Flow Parallel to Tube Bundles at Relative Pitches of 1.1 < s/d < 1.6," Heat Transfer-Soviet Research, (1975): 7, 3, pp. 84- 88.

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95. {{

}}2(a),(c) 96. Van den Eynde, G., "Comments on “A Resolution of the Stiffness Problem of Reactor Kinetics”," Nuclear Science and Engineering, (2006): 153:200-202. 97. Saha, P., and N. Zuber, "Point of Net Vapor Generation and Vapor Void Fraction in Subcooled Boiling," Proceedings Fifth International Heat Transfer Conference, September 3-7, 1974, Tokyo, Japan: 4:175-179. 98. Lahey, R.T., "A Mechanistic Subcooled Boiling Model," Proceedings Sixth International Heat Transfer Conference, August 7-11, 1978, Toronto, Canada: 1:293 - 297. 99. U.S. Nuclear Regulatory Commission, “An Integrated Structure and Scaling Methodology for Severe Accident Technical Issue Resolution,” NUREG/CR-5809, Appendix D, November 1991. 100. U.S. Nuclear Regulatory Commission, "Design-Specific Review Standard for NuScale SMR Design," Chapter 15, Section 15.6.6, Rev.0, June 2016. 101. U.S. Nuclear Regulatory Commission, NUREG-0800, Standard Review Plan for the Review of Safety Analysis Reports for Nuclear Power Plants: LWR Edition, Section 15.6.1, Revision 2, March 2007. 102. U.S. Nuclear Regulatory Commission, NUREG-0800, Standard Review Plan for the Review of Safety Analysis Reports for Nuclear Power Plants: LWR Edition, Section 15.0, Revision 2, March 2007. 103. American Nuclear Society, “Nuclear Safety Criteria for the Design of Stationary Pressurized Water Reactor Plants,” ANSI N18.2-1973. 104. U.S. Nuclear Regulatory Commission, NUREG-0800, Standard Review Plan for the Review of Safety Analysis Reports for Nuclear Power Plants: LWR Edition, Section 5.2.3, Revision 2, March 2007. 105. GE Nuclear Energy, “ABWR Design Control Document,” Revision 4, March 1997. 106. U.S. Nuclear Regulatory Commission, “Applying Statistics,” NUREG-1475, Rev. 1, March 2011. 107. U.S. Nuclear Regulatory Commission, "Design-Specific Review Standard for NuScale SMR Design," Chapter 15, Section 15.0, Rev. 0. June 2016. 108. U.S. Nuclear Regulatory Commission, NUREG-1503, "Final Safety Evaluation Report Related to the Certification of the Advanced Boiling Water Reactor Design, Main Report," July 1994 109. NuScale Power, LLC, "Technical Report: Containment Response Analysis Methodology," TR-0516-49084, Rev. 3. 110. NuScale Power, LLC, Topical Report "Non-LOCA Transient Analysis Methodology," TR-0516-49416-P-A, Rev. 3.

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Appendix A. Input for NuScale Power Module Loss-of-Coolant Accident Model

The purpose of this Appendix is to present inputs for initial and boundary conditions and other key parameters for the NPM LOCA input model. The core input parameters, plant initial conditions (i.e., operational conditions), safety-related system setpoints and delays, and the LOCA break spectrum parameters used for the calculations stated in this report are presented. The operational range is provided for several parameters along with the basis for selection of the upper or lower range value that conservatively bounds the LOCA response.

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A.1 Core Input Parameters

Table A-1. Core input parameters

{{

}}2(a),(c), ECI

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{{

}}2(a),(c) ECI

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A.2 Initial Plant Conditions

Table A-2. Initial conditions for loss-of-coolant accident analysis

{{

}}2(a),(c), ECI

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A.3 Safety-Related System Actuation Setpoints and Delays

Table A-3. Safety signal actuation setpoints and delays

{{

}}2(a),(c), ECI

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{{

}}2(a),(c), ECI

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A.4 Break Spectrum Parameters

Table A-4. Break spectrum parameters {{

}}2(a),(c), ECI

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Appendix B. Evaluation Model for Inadvertent Opening of RPV Valves

B.1 Introduction

B.1.1 Purpose

The purpose of this Appendix is to describe the evaluation model and methodology applied by NuScale to analyze NRC Standard Review Plan (Reference 101) event 15.6.1, Inadvertent Opening of a PWR Pressurizer Pressure Relief Valve, as well as the Inadvertent Operation of the Emergency Core Cooling System (ECCS) event as defined in Section 15.6.6 of the Design-Specific Review Standard (DSRS) for NuScale SMR Design (Reference 100). The methodology and EM are developed by extending the LOCA Methodology presented in the main body of this report. The methodology and EM follow the guidance provided in “Transient and Accident Analysis Methods,” Regulatory Guide (RG) 1.203 (Reference 1), and allow for demonstration via analysis that the acceptance criteria for Anticipated Operational Occurrences (AOOs) listed in DSRS Section 15.0 (Reference 107) are met.

For editorial convenience the 15.6.1 and 15.6.6 event scenarios may hereafter be collectively referred to as “IORV”, the Inadvertent Opening of an RPV Valve.

B.1.2 Scope

This appendix summarizes the following:

• Regulatory Requirements and Classification of the inadvertent RSV opening event (15.6.1) and inadvertent ECCS operation event (15.6.6) as AOOs • NPM design features important to the IORV event scenarios • Development of the 95/95 MCHFR limit for IORV Analysis • Description of the IORV NRELAP5 evaluation model, and changes from the LOCA EM. • Applicability for IORV analysis of NRELAP5 assessments against separate effects tests (SETs) and integral effects tests (IETs) • Applicability evaluation determining the adequacy of NRELAP5 for NPM IORV analyses • IORV analysis results and sensitivity studies

B.2 Background

The Inadvertent Opening of an RSV and the Inadvertent Operation of ECCS events exhibit a transient progression that is more similar to LOCAs than it is to other AOO events analyzed for the NPM. This progression is divided into two phases, similar to LOCA:

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• The first phase is initiated by the inadvertent opening of an RPV valve (RSV, RVV or RRV) that results in a blowdown of the RCS into the containment vessel (CNV). This scenario can be characterized as a steam region discharge (i.e., opening of an RSV or RVV) or a liquid region discharge (i.e., opening of an RRV). This phase ends when the ECCS valves open; when the event initiator is the inadvertent opening of an RVV or RRV, this stage ends when the remaining functioning ECCS valves open.

• The second phase begins when the ECCS valves open and ends when the NPM reaches a semi-equilibrium recirculating ECCS mode which defines the transition to long-term ECCS cooling.

These two phases align with Phase 1a and Phase 1b of the LOCA transient progression for the NPM as discussed in Section 9.1. The LOCA evaluation model has:

• identified and ranked important phenomena which occur during these transient phases for the NPM (Section 4.0),

• assessed NRELAP5 against separate effects tests and integral effects tests related to these phenomena (Section 7.0),

• determined NRELAP5 to be applicable for evaluating these phenomena (Section 8.0),

• and developed a conservative NRELAP5 LOCA analytical model for transient analyses which involve an un-isolatable decrease in RCS inventory event (Section 5.0).

Because of the phenomenological similarities to the LOCA pipe break events, a modified version of the LOCA evaluation model is used to analyze the IORV events for the NPM.

B.3 Regulatory Requirements for the Inadvertent Opening of an RPV Valve

The relevant requirements for NuScale design basis events are contained in the following Commission regulations:

• 10 CFR 20, “Standards for Protection Against Radiation”

• 10 CFR 50, “Domestic Licensing of Production and Utilization Facilities” (especially 10 CFR 50.46 and the general design criteria (GDC) of Appendix A)

• 10 CFR 100, “Reactor Site Criteria”

• 10 CFR 52, “Early Site Permits; Standard Design Certification; and Combined Licenses for Nuclear Power Plants”

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The Commission has provided guidance to meet these requirements in the NUREG- 0800 Standard Review Plan and a Design-Specific Review Standard (DSRS) for the NuScale SMR Design.

B.3.1 Classification of the NuScale 15.6.6 Inadvertent ECCS Event Scenario

In conventional PWRs the inadvertent ECCS actuation scenario typically results in an RCS inventory addition via actuation of the high pressure safety injection system (HPSI). Other protection systems such as containment isolation and the reactor protection system may be actuated as a result of the ECCS actuation. In the NuScale design however, inadvertent ECCS actuation results in a reduction of RPV liquid inventory via the opening of an ECCS valve (RVV or RRV). No other protection systems are actuated because in the NuScale design the ECCS operates independently of other systems. Although the inadvertent opening of an ECCS valve results in a reduction of RPV inventory, the core remains covered with coolant and any coolant released through the open ECCS valve is retained and cooled by the steel containment, and is eventually recirculated back to the RPV via the RRVs.

While the inadvertent ECCS event in the NuScale design is thermal-hydraulically similar to the loss of coolant accident scenario, the inadvertent ECCS event is more properly categorized as an AOO.

The AOO acceptance criterion iii (discussed in Section B.3.3) generates two key questions regarding the inadvertent ECCS actuation event for the NuScale design:

1. Does ECCS actuation in response to an AOO, inadvertent opening of a single valve, or inadvertent ECCS actuation generate a postulated accident?

2. Does ECCS actuation in response to an AOO, inadvertent opening of a single valve, or inadvertent ECCS actuation result in a consequential loss of function of the RCS barrier?

The AOO acceptance criterion iii in the DSRS states that it is based on ANS standards. Additional information about the event classification and associated ANS standards are provided in the SRP (Reference 102). The SRP states that “Postulated accidents are also known as Condition IV events in the unofficial ANS standards.” Per the definition of Condition IV events in ANSI N18.2 (Reference 103), postulated accidents are “not expected to occur, but are postulated because their consequences would include the potential for the release of significant amounts of radioactive material.” The examples of Condition IV events cited in ANSI N18.2 involve significant component failures which “would include the potential for the release of significant amounts of radioactive material”, e.g., “major rupture of a pipe containing reactor coolant up to and including double-ended rupture of the largest pipe in the reactor coolant pressure boundary.”

In the NuScale design, ECCS actuation in response to an AOO, inadvertent opening of a single valve, or inadvertent ECCS actuation do not in themselves present a potential for release of significant amounts of radioactivity. All reactor coolant released from the RPV

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is captured by the CNV, cooled, and eventually circulated back to the RPV. The core remains covered with liquid coolant at all times. Therefore, based on a review of SRP 15.0, these NuScale events do not generate a postulated accident and thus may be conservatively categorized as AOOs, which have more restrictive acceptance criteria than accidents.

Regarding the criterion that an AOO must not result in the “consequential loss of function of the RCS or reactor containment barriers,” the SRP also applies this acceptance criterion to ANSI N18.2 Condition II (“Incidents of Moderate Frequency”) and Condition III (“Infrequent Incidents”) events. ANSI N18.2 presents the following scenarios as examples of Condition II and III events:

Condition II: “depressurization by spurious operation of an active element, for example, relief valve, pressurizer spray valve.”

Condition III: “loss of reactor coolant, such as from a small ruptured pipe or from a crack in a large pipe, which would prevent orderly reactor shutdown and cooldown assuming makeup is provided by normal makeup systems only.” (i.e., small-break LOCA)

Thus, by definition, Condition II and III events do not in themselves result in a consequential (significant) loss of function of the RCS barrier. The two conventional reactor examples given above are characterized by continuous release of reactor coolant from the RCS, either through a valve (Condition II) or through a small-break LOCA (Condition III), and neither results in a consequential loss of function of the RCS or containment barriers. This continuous release (and recirculation) of reactor coolant without consequential loss of function of the RCS or containment barriers is exactly what occurs in the NuScale design for the ECCS actuation in response to an AOO, inadvertent opening of a single valve, or inadvertent ECCS actuation event scenarios.

SRP Section 5.2.3 (Reference 104) describes a gross failure of the reactor coolant pressure boundary as a “substantial reduction in capability to contain reactor coolant inventory, reduction in capability to confine fission products, or interference with core cooling”. Therefore a “substantial loss of function” requires a gross failure of the RCS barrier. This is because the function of fission product confinement is integrated with the functions of inventory control and heat removal, i.e., the function of fission product confinement is maintained if the functions of inventory control and heat removal are maintained. The requirement to maintain the “fission product barrier” does not mean that leakage from fuel defects, or activation products in RCS coolant must be confined in the RCS following all events within the design basis. It refers to maintaining the integrity of the cladding. Without fuel cladding failure there are no significant radiological consequences associated with an event, and therefore there can be no “consequential loss of function” of the RCS barrier.

Based on a review of DSRS 15.0 and SRP 15.0, the ECCS actuation in response to an AOO, inadvertent opening of a single valve, or inadvertent ECCS actuation events do not result in a substantial reduction in capability to contain reactor coolant inventory,

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reduction in capability to confine fission products, or interference with core cooling. Therefore, these events do not result in a consequential loss of function of the RCS barrier and thus may be conservatively categorized as AOOs.

B.3.2 Design-Specific Review Standard Definition of AOOs

The following discussion of categorization of transients and accidents is excerpted from Section 15.0 of the NuScale DSRS (Reference 107):

“Categorization According to Frequency of Occurrence. Each initiating event is categorized as either an anticipated operational occurrence (AOO); a postulated accident, which includes the infrequent event (IE) classification; or special event.

AOOs, as defined in Appendix A to 10 CFR Part 50, are those conditions of normal operation that are expected to occur one or more times during the life of the nuclear power unit.

….

Postulated accidents and infrequent events are unanticipated occurrences that are postulated but not expected to occur during the life of the nuclear power unit.”

B.3.3 AOO Acceptance Criteria

Section 15.0 of the NuScale DSRS (Reference 102) lists the analysis acceptance criteria for AOOs that are necessary to meet the relevant regulatory requirements:

i. Pressure in the reactor coolant and main steam systems should be maintained below 110 percent of the design values in accordance with the American Society of Mechanical Engineers (ASME) Boiler and Pressure Vessel Code.

ii. Fuel cladding integrity shall be maintained by ensuring that the minimum departure from nucleate boiling ratio (DNBR) remains above the 95/95 DNBR limit.

The reviewer applies a third criterion, based on the ANS standards to ensure that there is no possibility of initiating a postulated accident with the frequency of occurrence of an AOO.

iii. An AOO should not generate a postulated accident without other faults occurring independently or result in a consequential loss of function of the RCS or reactor containment barriers.

The AOO acceptance criteria for the IORV analysis are provided in Table B-1.

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Table B-1. Anticipated Operational Occurrence Regulatory Acceptance Criteria

Condition Description Criterion

Reactor Coolant System Peak Pressure ≤ 110% of Design

Secondary System Peak Pressure(1) ≤ 110% of Design

Minimum Critical Heat Flux Ratio ≥ Limit

Maximum Fuel Centerline Temperature ≤ UO2 melting temp

Generate More Serious Plant Condition? No

Containment Pressure(2) ≤ 100% of Design

(1) The “secondary system” refers to the region between the FWIVs and the MSIVs.

(2) The containment pressure response is evaluated in a separate analysis that bounds all other events, including the IORV event.

B.3.4 Regulatory Guidance on Analysis Assumptions

The NuScale DSRS (Reference 100 Section 15.6.6 Subsection II) lists four assumptions regarding important parameters that shall be considered. These are also applicable to the 15.6.1 event:

“The initial power level is taken as the licensed core thermal power for the number of loops initially assumed to operate plus an allowance of 2 percent to account for power measurement uncertainties unless the applicant can justify a lower power level. The operating condition at the initiation of the event should correspond to the operating condition that maximizes the consequences of the event.”

“Applicant should conservatively assume the maximum time delay and the most reactive rod held out of the core.”

“The core burnup is selected to yield the most limiting combination of moderator temperature coefficient, void coefficient, Doppler coefficient, axial power profile, and radial power distribution.”

“Mitigating systems should be assumed to be actuated in the analyses at setpoints with allowance for instrument inaccuracy in accordance with RG 1.105.”

All of these assumptions are applicable to the IORV analysis.

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B.3.5 SRP Section 15.6.1 “Inadvertent Opening of a PWR Pressurizer Pressure Relief Valve or a BWR Pressure Relief Valve”

The discussion in SRP 15.6.1 (Reference 101), which characterizes inadvertent RCS valve openings in current generation plants, has relevance to the NuScale design. Accidental depressurization of the RCS through a pressure relief valve (PORV or code safety valve) is generally categorized as an AOO even though the rate of loss of coolant exceeds the makeup capacity:

“An accidental depressurization of the reactor coolant system (RCS) could be caused by the inadvertent opening of a pressure relief valve, which in turn could be caused by a spurious electrical signal or by an operator error. As this event can occur one or more times during the plant’s lifetime, it is an anticipated operational occurrence (AOO), as defined in 10 CFR Part 50, Appendix A.”

Thus, based on the guidance in SRP 15.6.1, classifying the inadvertent opening of a single ECCS valve in the NuScale design as an AOO is consistent with SRP 15.6.1, even if such an event would exceed the makeup capacity of the CVCS.

B.3.6 10 CFR 50.46 Considerations

10 CFR 50.46 (Reference 3) defines LOCAs as follows:

“(c) As used in this section: (1) Loss-of-coolant accidents (LOCA's) are hypothetical accidents that would result from the loss of reactor coolant, at a rate in excess of the capability of the reactor coolant makeup system, from breaks in pipes in the reactor coolant pressure boundary up to and including a break equivalent in size to the double-ended rupture of the largest pipe in the reactor coolant system.”

A literal interpretation of 10 CFR 50.46 concludes that LOCAs are pipe breaks in the RCS pressure boundary that result in coolant loss rate beyond the ability of the CVCS (makeup) to compensate for the break flow. Per Section B.3.1 of this report, the SRP defines accidents as events that are never expected to occur during the plant lifetime, and “are postulated because their consequences would include the potential for the release of significant amounts of radioactive material.”

Applying these definitions for loss of coolant accidents to the NuScale design, an ECCS actuation in response to an AOO, an inadvertent opening of a single valve, or an inadvertent ECCS actuation does not meet either the letter (or the intent) of the LOCA regulations. Operation of the ECCS valves is not a pipe break, it is a normal plant response for certain AOOs and accident scenarios, and in itself does not pose a potential for the release of significant amounts of radioactive material. This categorization of inadvertent RPV valve openings not being a LOCA is consistent with the defined scope of the LOCA EM as defined in Section 1.2 of this report. However, the scope notes that the LOCA EM may be used to evaluate such transients.

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B.3.7 Regulatory Precedent

Safety relief valves (SRVs) are used in BWR designs for heat removal by venting reactor coolant to the suppression pool for mitigation of AOOs. The SRVs pneumatically open after receiving an automatic or manual actuation signal (functioning as pressure relief valves) or they self-actuate from inlet steam pressure (functioning as code safety valve). In the ABWR DCD (p5.2-1, Reference 105) the safety design bases of the pressure relief system (SRVs) are:

• “Prevent overpressurization of the nuclear system that could lead to the failure of the RCPB.” • “Provide automatic depressurization for small breaks in the nuclear system occurring with maloperation of both the RCIC System and the HPCF System so that the low pressure flooder (LPFL) mode of the RHR System can operate to protect the fuel barrier.”

The power generation design bases of the SRVs are (p5.2-1, Reference 105):

• “Discharge to the containment suppression pool.” • “Correctly reclose following operation so that maximum operational continuity is obtained.”

The final bullet above confirms that reclosing the SRVs after discharging reactor coolant to the suppression pool is not a safety function in the ABWR design. In other words, closure of these valves is not required to prevent loss of reactor coolant or the release of radioactivity material entrained in coolant to the suppression pool, but is required only to support a return to power operation.

The ABWR, which the NRC issued a Final Safety Evaluation Report (Reference 108) for, provides precedent that opening SRVs does not result in a reduction in capability to confine fission products (which is a function of the RCS barrier). This is illustrated by considering that the ABWR inadvertent MSIV closure event is analyzed as an AOO. For analysis of this event, the RCS depressurization rate and resultant blowdown to containment is maximized to demonstrate that 10 CFR Part 20 limits are met even when the reactor has been operating with defective fuel (p 15.2-16, Reference 105).

Consistent with the ABWR precedent, reclosing the ECCS valves in the NuScale design is not a safety function, but is a required step before returning the module to power operation. When opening the NuScale ECCS valves in response to an AOO, fission products are confined to the RCS because the fuel cladding barrier is not compromised. Similar to the ABWR precedent it can be shown that 10 CFR Part 20 limits are met when actuating ECCS in response to an AOO for the NuScale design.

B.3.8 IORV Classification Conclusion

The following statements are true based on the foregoing discussions:

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• Both inadvertent opening of a single ECCS valve and inadvertent ECCS actuation are properly categorized as AOO events. Inadvertent ECCS actuation resulting in all valves opening at full pressure is considered a beyond-design-basis event because the ECCS valve design includes an inadvertent actuation block mechanism (see Section B.4.1). • Neither inadvertent opening of a single ECCS valve nor inadvertent ECCS actuation result in a LOCA in the context of 10 CFR 50.46. • ECCS actuation in response to an AOO is allowable within the NRC’s regulatory framework, specifically considering Standard Review Plan (SRP) Section 15.0.

B.4 NuScale Design Considerations for IORV Events

The NuScale ECCS and IAB are described in Section 3.3.1 of this report.

B.4.1 Inadvertent ECCS Signal

An inadvertent ECCS signal could signal for opening of all five ECCS valves, or one division of ECCS (1 RRV and 2 RVV).

The ECCS valve design includes an inadvertent actuation block (IAB) feature that prevents valve opening until the RCS has depressurized below a set differential pressure with the containment. The module protection system (MPS) ECCS actuation signals are unique to actuating ECCS (ECCS actuation does not initiate a reactor trip or other engineered safety features). If an inadvertent ECCS actuation signal is generated at normal operating pressure, the IAB trip solenoid valves will open and a small amount of liquid in these lines will be released into containment; however, this is not expected to be sufficient to generate a high containment pressure signal. Therefore, if an inadvertent ECCS actuation signal is generated, the ECCS valves will not open due to the IAB and no other engineered safety feature will be actuated. It is expected that the operators will take action to address the inadvertent signal without resulting in ECCS valve opening or other actuation of the module protection system. Therefore, an inadvertent ECCS signal is not a latent condition that need be considered as part of other design basis events.

B.4.2 Inadvertent opening of a single ECCS valve

A mechanical failure could result in the opening of a single ECCS valve. The valve opening is considered the initiating event (i.e. the valve opening is not the result of an initiating event plus a coincident single failure).

With respect to the event frequency and classification, with the inadvertent actuation block device, the inadvertent opening of a single ECCS valve is not expected to occur during the lifetime of the plant; the point estimate is on the order of 10-5 per module year; for a 60 year operating life, this corresponds to a frequency of about 6x10-4 events over the life of the module).

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Although an ECCS valve is not expected to inadvertently open during the life of a module, the event is conservatively categorized as an AOO and analyzed against the AOO acceptance criteria.

B.4.3 Inadvertent opening of multiple ECCS valves

As discussed in Section B.4.1 the initiating event of an inadvertent ECCS actuation signal while the RCS is at normal pressure conditions does not result in ECCS valve opening without other failures. Simultaneous mechanical failures on multiple ECCS valves are beyond design basis with respect to identifying initiating events.

Therefore, with the ECCS valve IAB feature, simultaneous inadvertent opening of multiple ECCS valves at normal RCS pressure is considered beyond the design basis.

B.4.4 Inadvertent opening of one reactor safety valve

The reactor safety valves (located on top of the pressurizer) provide over-pressure protection of the reactor pressure vessel. Mechanical failure associated with the valve internals could result in the inadvertent opening of a reactor safety valve.

Consistent with the categorization of the inadvertent opening of one ECCS valve, this event is categorized as an AOO and analyzed against AOO acceptance criteria. This is consistent with the SRP (Reference 102, p15.0-2).

B.5 MCHFR Limit for IORV Analysis

NRELAP5 heat transfer {{

}}2(a),(c),ECI

B.5.1 KATHY CHF Tests

The NuFuel HTP2™ Critical Heat Flux test program conducted at the KATHY Laboratories in Karlstein, Germany included tests to characterize thermal mixing and steady and transient CHF performance across a range of thermal-hydraulic conditions. The various tests made use of a {{ }}2(a),(c),ECI KATHY testing involved four separate configurations including:

• {{

}}2(a),(c),ECI

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• {{

}}2(a),(c),ECI

The tests were performed in the pressurized water reactor loop shown in Figure B-1. The axial test section is provided in Figure B-2 for the uniform profile cases. The section includes a channel box (flow channel), fuel simulators, spacer grids (HMP, HTP, and SSG), and instrumentation. The portion of the test section containing the actively heated portion of the fuel simulators was modeled with NRELAP5. In Figure B-2, this is the region from the Bottom of the Heated Length (BOHL) to the End of the Heated Length (EOHL). The test section radial layouts are shown in Figure B-3.

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{{

}}2(a),(c),ECI

Figure B-1. KATHY Test Loop

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{{

}}2(a),(c),ECI

Figure B-2. KATHY NuFuel HTP2TM Test Section Axial Layout

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{{

}}2(a),(c),ECI

Figure B-3. KATHY {{ }}2(a),(c),ECI Radial Layouts

The KATHY CHF tests were performed in the following manner:

• {{

}}2(a),(c),ECI

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− {{ }}2(a),(c),ECI

The inlet conditions for the KATHY tests used for the CHF correlation are shown in Table B-2 below:

Table B-2. KATHY Inlet Boundary Condition Ranges

{{

}}2(a),(c),ECI

The inlet mass flux, temperature, and pressure ranges for the KATHY tests bound the NPM core inlet conditions for the IORV event scenarios at the time of minimum CHFR. Therefore it is appropriate to utilize the minimum CHFR design limit derived from the KATHY CHF data. The KATHY data have overlap with Stern Laboratories data (Section 7.3) by design. Comparisons were made between Stern and KATHY by plotting CHF versus mass flux for a particular pressure and inlet subcooling level. There is direct overlap at {{

}}2(a),(c),ECI used for the KATHY tests differs from that used in the Stern tests.

B.5.2 NRELAP5 Model of KATHY Test Section

The KATHY NuFuel HTP2™ test section is modeled using a {{

}}2(a),(c)

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{{

}}2(a),(c)

B.5.3 CHF Correlation Limit

A total of {{

}}2(a),(c) data and applies to the NuFuel-HTP2™ fuel design. The methodology for determining the correlation limit is identical to that used in the development of the NSP2 and NSP4 CHF correlations and is illustrated by Figure B-4.

{{

}}2(a),(c)

Figure B-4. CHF Statistical Methods Flow Chart

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Normality of samples was tested with either {{ }}2(a),(c) These tests are described in Reference 106.

Variance between samples was tested with either {{

}}2(a),(c)

Subsets of data were identified and binned based on test ID, pressure, mass flux, inlet subcooling, and exit quality. Using the statistical methods illustrated in Figure B-4 these subsets were combined into the composite subsets. Tolerance limits for each composite subset were calculated, with the results indicating that a CHFR limit of 1.05 is sufficient to guarantee CHF will not occur with 95% probability and 95% confidence. The maximum limiting value for the KATHY NuFuel HTP2™ data sets as reported by the Hench Levy correlation is 1.12.

Additional penalties are applied to the 95/95 1.05 design limit consistent with subchannel methodology: a 3% engineering uncertainty factor and a 3% fuel rod bowing factor. Therefore, the safety limit applied for inadvertent RPV opening analysis is 1.05 * 1.03 * 1.03 = 1.114.

Since no KATHY NuFuel HTP2™ CHF test data exists for the modified Griffith-Zuber low flow condition (<0.1 Mlbm/ft2-hr), the 1.29 CHF limit derived from the Stern CHF data in

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Section 7.3.6 is used for the IORV EM. The same additional penalties for 3% engineering uncertainty and 3% rod bowing are applied, resulting in a final CHFR limit of 1.37 for low flow conditions. For IORV analyses the CHF limit is typically challenged within the range of the high flow correlation, and an adequately large margin exists when flow transitions to the low flow CHF correlation range.

{{

}}2(a),(c),ECI

Figure B-5. HS171 Correlation: Predicted vs Measured Power

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B.6 IORV NRELAP5 Evaluation Model Description

The NRELAP5 model utilized for IORV analysis is similar in all important aspects to the LOCA Evaluation Model described in Section 5.1 of this report, and is derived from an updated IORV model that is based on the current NPM NRELAP5 plant model which incorporates the latest NPM design specifications. The NRELAP5 LOCA modeling methods are used for analysis of IORV events because in the NPM design the transient phenomena of these events are similar to the LOCA pipe break events. Certain modifications of the LOCA EM methodology are applied to better align the analysis with the AOO acceptance criteria instead of the accident criteria. These modifications are shown in Table B-3 below:

Table B-3. Changes to LOCA EM for IORV EM

Parameter / LOCA EM IORV EM Rationale Component Reactor Safety Valves The minimum flow model is RSVs modeled to applied for the RSVs because RSV flow capacity produce minimum Same as LOCA EM any larger RSV flows are required flow. bounded by the inadvertent RVV opening event. 0.1 sec RSV stroke time is RSV opening 1.0 sec 0.1 sec consistent with RVV opening stroke time stroke time. Moody/Henry-Fauske critical Inadvertently opened flow model maximizes two- valve uses Moody/Henry- RSV critical flow Henry-Fauske phase flow, and maintains Fauske (c=3). Remaining model (c=2) consistency with LOCA break RSV uses Henry-Fauske methodology described in (c=2). Section 6.6.1 of this report. ECCS Valves (RRVs and RRVs) Maximum flow capacity maximizes RPV Flow areas and flow ECCS valve flow depressurization rate, which is coefficients set to the Same as LOCA EM capacity conservative for MCHFR. This largest design values. is confirmed via minimum ECCS sizing sensitivity cases. Faster opening produces ECCS valve higher flow rates and faster opening stroke 4.0 sec 0.1 sec depressurization, which is time conservative for MCHFR. Moody/Henry-Fauske critical Inadvertently opened flow model maximizes two- valve uses Moody/Henry- ECCS valve Henry-Fauske phase flow, and maintains Fauske (c=3). Remaining critical flow model (c=2) consistency with LOCA break ECCS valves use Henry- methodology described in Fauske (c=2). Section 6.6.1 of this report.

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Parameter / LOCA EM IORV EM Rationale Component Reactor Core and Kinetics The core burnup is selected to yield the most limiting combination of BOC kinetic parameters moderator and Doppler are considered bounding, reactivity feedback, axial with the smallest prompt power profile, and radial neutron lifetime applied to power distribution. maximize core initial The prompt neutron Consistent with AOO Fuel burnup and energy by prolonging the lifetime and the effective Regulatory Analysis kinetics fission power transient. delayed neutron fraction Assumptions (Section B.3.4). Biasing is applied to all (βeffective) are applied kinetic parameters to consistent with the account for uncertainty in assumed fuel burnup. calculated values Biasing of kinetics (Section 5.1.2.2.5). parameters for uncertainty is identical to LOCA EM. Minimum bounding scram Minimum bounding scram worth consistent with worth over any time in Consistent with AOO assumed time in fuel Scram worth fuel cycle is applied. Most Regulatory Analysis cycle is applied. Most reactive control rod not Assumptions (Section B.3.4). reactive control rod not credited. credited. Applies conservative signal actuation delay Consistent with AOO times from the time the Same signal actuation Regulatory Analysis process setpoint is delays as LOCA EM, but Scram delay Assumptions (Section B.3.4), reached to the time the additional 2-sec scram and analysis of other NPM rod start to fall. An delay is not applied. AOO events. additional 2 sec delay is added for conservatism. Reactivity feedback “Least negative” reactivity parameters (with Reactivity feedback feedback maximizes transient uncertainties) are chosen parameters (with fission power, which is conservatively by using uncertainties) are applied Reactivity conservative for MCHFR. the least negative to achieve the most feedback feedback coefficients to limiting (least negative) Consistent with AOO maximize the energy feedback for the assumed Regulatory Analysis deposition due to fission time in fuel cycle. Assumptions (Section B.3.4). power. (Section 5.1.2.2.5) A 15% bias is included in the NRELAP5 UO2 thermal conductivity and The 15% bias is not Fuel Thermal Consistent with other AOO heat capacity tables to applied to the UO2 Properties events. increase stored thermal thermal properties. energy (Section 5.1.2.2.4).

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Parameter / LOCA EM IORV EM Rationale Component A minimum gap conductance A minimum (BOL) fuel- maximizes the stored thermal cladding gap Gap conductance is energy in the fuel. However a conductance varied via sensitivity maximum gap conductance Gap Conductance conservatively bounds analyses using bounding allows the fuel thermal energy and maximizes the initial maximum and minimum to be released to the cladding stored energy in the fuel values. at a higher rate. Either (Section 5.1.2.2.4). assumption could produce a limiting CHFR. A bounding axial power A single limiting generic shape (over all fuel Use of subchannel analysis axial power shape is used cycles) from the steady- axial shapes is consistent with for both of the core state subchannel analysis other AOO analyses. channels. Sensitivity is applied for all base calculations (Section cases. The same axial Consistent with AOO 9.6.6) using generic top- Axial power shape power shape is applied to Regulatory Analysis peaked, bottom-peaked, both the hot channel and Assumptions (Section B.3.4), and chopped cosine the average channel. sensitivity analyses are profiles show that axial Sensitivity cases are performed to determine which power shape has performed to confirm that axial shape is most limiting for negligible impact on the selected axial shape MCHFR. LOCA FOMs. is conservatively limiting. The radial power peaking is A limiting assembly radial consistent with other AOO Hot assembly peaking 1.4 peaking factor derived analyses. is selected to bound all using subchannel possible power peaking. methodology is applied to The radial peaking, in Radial Power Sensitivity study the hot channel. Biases combination with the Peaking performed with different for measurement NRELAP5 CHF correlation, is radial hot assembly uncertainty and used to demonstrate peaking values. (Table engineering uncertainty compliance with the AOO A-1)”. are also applied. 95/95 DNBR acceptance criterion (Section B.3.3).

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Parameter / LOCA EM IORV EM Rationale Component CHF is calculated using CHF calculated using For the 15.6.1 and 15.6.6 NRELAP5 heat transfer NRELAP5 heat transfer event scenarios, analysis option 171, which applies option 171, same as shows that MCHFR always the extended Hench-Levy LOCA EM. occurs early in the transient CHF correlation that uses before reactor trip when the a heat balance approach The 100% data coverage core flow is in the high mass with pressure correction CHFR limit for the KATHY flux range (≥135.6 kg/m2-s) for high flow conditions, CHF data sets is 1.12, where the Extended Hench- and the Griffith-Zuber with a 95/95 tolerance Levy CHF correlation is CHF correlation with high limit of 1.05. Applying a applied. The KATHY NuFuel void correction term for 3% engineering HTP2™ CHF tests were low flow conditions. uncertainty factor and 3% performed by AREVA with uncertainty for fuel rod heated assembly A CHFR analysis limit of bow produces a MCHFR configurations that are very 1.29 was determined to 95/95 tolerance limit of representative of the NuScale envelope the Stern CHF 1.114. A conservative limit fuel design. The inlet mass MCHFR Limit test data for both of 1.13 is applied in the flux, temperature, and pressure correlations (see Sections IORV analyses when in ranges for the KATHY tests 2.2.2 and 7.3.6). the high mass flux range bound the NPM core inlet (≥135.6 kg/m2-s). conditions for the 15.6.1 and 15.6.6 event scenarios at the For low flow conditions, time of MCHFR. A total of 597 the Stern CHF limit (1.29) KATHY CHF data points were with 3% engineering examined with NRELAP5. uncertainty factor and 3% uncertainty for fuel rod bow is applied. This produces a MCHFR 95/95 tolerance limit of 1.37 when in the low mass flux range (<135.6 kg/m2-s).

B.6.1 Electric Power Availability

The analysis assumptions regarding electric power availability for the inadvertent RPV opening event are the same as for the LOCA methodology: • All electric power is available • Loss of normal alternating current (AC) power • Loss of normal AC and DC power

Sensitivity cases are run to determine which electric power availability condition is the most limiting for the inadvertent RPV opening acceptance criteria.

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B.6.2 Single Active Failure Evaluation

The analysis assumptions for single active failure for the inadvertent RPV opening event are the same as for the LOCA methodology: • No single failure • Failure of a single RVV to open • Failure of a single RRV to open • Failure of one ECCS division (i.e., one RVV and one RRV)

Sensitivity cases are run to determine which single failure scenario is the most limiting for the inadvertent RPV opening acceptance criteria.

B.6.3 Initial Conditions and Biasing

Table B-4 shows a comparison of the initial conditions applied in the IORV analysis with those from the LOCA analysis in Tables 5-6 and A-2.

Table B-4. Comparison of LOCA and IORV initial conditions

LOCA IORV Process Range Range Basis Parameter (Nominal) (Nominal) Initial core power increased by 2 percent to account for measurement uncertainty, consistent Core power 102 percent 102 percent with IORV regulatory analysis requirements (Section B.3.4) For IORV, high RCS temperature places the RCS closer to saturated conditions at the time of transient initiation. Low RCS temperature RCS average 535–555 °F 535–555 °F reduces level swell during the blowdown, and temperature (545 °F) (545 °F) delays the time until a two-phase choking condition exists at the RVV or RSV. Two-phase choking restricts flow through the valves and limits the rate of depressurization. For IORV, high initial pressure maximizes initial RCS energy. Low initial pressure places the RCS Pressurizer 1780–1920psia 1780–1920psia closer to saturation at transient initiation. Either pressure (1850 psia) (1850 psia) extreme can affect the timing of two-phase flow through the inadvertently opened valve. For IORV, high initial level minimizes the pressurizer steam volume, causing faster RCS depressurization following RSV or RVV inadvertent opening. High initial level also results 52–68 percent 52–68 percent in earlier two-phase choking due to level swell, Pressurizer level (60 percent) (60 percent) reducing flow through an inadvertently opened RVV or RSV. Low initial level increases the initial steam volume, slowing the depressurization rate, and also delays the time to two-phase choking through the RVV or RSV.

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LOCA IORV Process Range Range Basis Parameter (Nominal) (Nominal) LOCA applies maximum 2 psia to maximize CNV peak pressure (Table A-2).

Containment 0.037 – 2.0psia 0.037 – 2.0psia For IORV, the early timing of MCHFR causes the pressure (0.037 psia) (0.037 psia) initial CNV pressure to have no effect on MCHFR results. Any initial CNV pressure (within the operating range) will result in choked flow through the inadvertently opened valve. Consistent with LOCA (Table A-2), IORV Main steam initializes with maximum 535 psia to maximize 465 – 535psia 465 – 535psia pressure at 100 overall system energy. The high initial pressure (500 psia) (500 psia) percent power bias is also necessary to satisfy the AOO peak pressure acceptance criterion (Section B.3.3). Consistent with LOCA (Table A-2), IORV applies Feedwater maximum 310 °F to maximize overall system temperature at 290–310 °F 290–310 °F energy. The high initial FW temperature bias is 100 percent (300 °F) (300 °F) also necessary to satisfy the AOO peak pressure power acceptance criterion (Section B.3.3) Sensitivity cases on RCS flow are run for IORV. Low flow minimizes steady state CHFR at RCS flow at 100 transient initiation. High flow reduces the 535–670 kg/s 535–670 kg/s percent power temperature difference across the core, resulting in an increased core inlet temperature when a fixed average temperature is used. IORV applies a target core bypass flow of 8.5% Bypass flow 8.5 percent of 8.5 percent of of the total system flow, consistent with the (reflector and ≈ ≈ subchannel analysis methodology, A tolerance guide tubes) total core flow total core flow band of 8-9% is considered acceptable. For IORV, the early timing of MCHFR causes the initial pool temperature to have no effect on Reactor pool 40–140 °F 40–140 °F MCHFR results. A high initial pool temperature temperature (100 °F) (100 °F) minimizes heat transfer to the pool and is conservative for evaluation of the AOO maximum pressure acceptance criterion (Section B.3.3). For IORV, the early timing of MCHFR causes the initial pool level to have no effect on MCHFR Reactor pool 55 – 68 ft 55 – 68 ft results. A low initial pool level minimizes heat level (68 ft) (68 ft) transfer to the pool and is conservative for evaluation of the AOO maximum pressure acceptance criterion (Section B.3.3).

B.7 NRELAP5 Assessments and Applicability

The NRELAP5 assessments discussed in Section 7.0 in support of the LOCA EM are also applicable for the IORV EM because of the physical phenomena and their importance ranking between the two scenarios are the same. Consequently, the bottom- up and top-down applicability evaluations presented in Sections 8.2 and 8.3 are also valid for the IORV EM. The following subsections review the applicability to the IORV EM of the previously documented SET and IET tests, and also provide additional NRELAP5 assessment results for NIST-1 test HP-43, an updated RVV spurious opening test.

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B.7.1 High-Ranked IORV Phenomena

For the NuScale NPM there are no significant differences in physics phenomena between the LOCA and IORV events. Although the initiating events are different (pipe break vs. inadvertent valve opening), the governing thermal hydraulic and core physics mechanisms are identical. Therefore the high-ranked phenomena from the LOCA PIRT shown in Table 4-4 also apply to the IORV event scenarios. The PIRT event phases of initial blowdown (1a) and ECCS actuation (1b) are also identical for LOCA and IORV.

B.7.2 Separate Effects Tests

B.7.2.1 Legacy Tests

The NRELAP5 assessments against the separate effects tests discussed in Section 7.2 also support the use of NRELAP5 for analysis of IORV events. Since the governing thermal hydraulic and core physics mechanisms are identical for the LOCA and IORV events, the legacy assessments discussed in Section 7.2 are also applicable for the IORV EM.

B.7.2.2 Stern Critical Heat Flux Tests

The assessment of the NRELAP5 CHF correlations via the NuScale Stern Critical Heat Flux Tests documented in Section 7.3 could also be applied to the IORV event scenario. However, for the IORV EM the data from the NuScale KATHY CHF tests is used to compute the 95/95 MCHFR limit required for AOO analysis. The KATHY tests and development of the 95/95 MCHFR limit is discussed in Section B.5.

B.7.2.3 SIET Steam Generator Tests

The NRELAP5 assessments against the experiments conducted under the NuScale testing program at SIET laboratories that are discussed in Section 7.4 also support the use of NRELAP5 for analysis of IORV events. Since the governing thermal hydraulic and mechanisms are identical for the LOCA and IORV events, the LOCA SIET test assessments are also applicable to the IORV EM.

B.7.3 NIST-1 Integral Effects Test Assessments

B.7.3.1 HP-09 RVV Opening Test

The NRELAP5 assessment for the NIST-1 HP-09 Spurious RVV opening test is discussed in Section 7.5.8 of this report. The assessment results (Section 7.5.8.5) show that:

• The RVV mass flow rate (Figure 7-117) is over-predicted by NRELAP5 during {{ }}2(a),(b),(c),ECI of the transient. Thereafter, the calculated flow shows excellent agreement with the measured flow rate.

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• The NRELAP5 calculated pressurizer pressure (Figure 7-118) shows excellent agreement with the data over the entire 6000 second duration of the test. An examination of the first {{ }}2(a),(b),(c),ECI of the pressurizer pressure (Figure 7-119) shows that the NRELAP5 predicted pressure is higher than the measured pressure. • The CNV pressure comparison over the {{ }}2(a),(b),(c),ECI is shown in Figure 7-121. The peak pressure from data and model are 531.3 psia and 545.5 psia, respectively. The comparison shows reasonable-to-excellent agreement with the measured data. • The pressurizer and RPV levels are compared in Figure 7-122 and Figure 7-123, respectively. The comparisons show reasonable-to-excellent agreement over the full duration of the test. The short term pressurizer level decrease over {{ }}2(a),(b),(c),ECI is under-predicted by NRELAP5, although this difference is not apparent in the short term RPV level comparison (Figure 7-124) which shows excellent agreement.

B.7.3.2 HP-43 Updated RVV Opening Test

The HP-43 test is an updated version of the HP-09 test, used to assess the capability of NRELAP5 to predict the integral response of the NIST-1 facility to {{

}}2(a),(c)

The important design and initial condition differences between the HP-43 and HP-09 tests are shown in Table B-5:

Table B-5. NIST-1 Spurious RVV Test Differences {{

}}2(a),(b),(c),ECI

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The Figures-of-Merit for the HP-43 test are: RPV pressure, RPV level, CNV pressure, and CNV level. The NIST-1 facility description, phenomena addressed, experimental procedure, and special analysis techniques are similar or identical to those described for the HP-09 assessment in Section 7.5.8.

B.7.3.3 Assessment Results

Figure B-6 through Figure B-11 present HP-43 transient short-term (0-800 seconds) code-to-data comparisons of selected parameters. The FOM comparisons of pressurizer pressure, RPV level, CNV pressure, and CNV level show reasonable-to-excellent agreement. It should be noted that ECCS actuation occurred at approximately 1190 seconds and does not appear in the short-term period plots.

The code-to-data comparison of pressurizer pressure presented in Figure B-6 shows reasonable-to-excellent agreement, with a slight under-prediction by NRELAP5 {{ }}2(a),(b),(c),ECI

{{

}}2(a),(b),(c),ECI

Figure B-6. HP-43 transient short-term pressurizer pressure comparison

The code-to-data comparison of pressurizer level presented in Figure B-7 shows reasonable-to-excellent agreement. Note that the {{

}}2(a),(b),(c),ECI

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{{

}}2(a),(b),(c),ECI

Figure B-7. HP-43 transient short-term pressurizer level code-to-data comparison

The code-to-data comparison of short-term RPV level presented in Figure B-8 shows reasonable-to-excellent agreement. It is notable that the timing of the minimum RPV level occurs {{ }}2(a),(b),(c),ECI in the NRELAP5 results.

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{{

}}2(a),(b),(c),ECI

Figure B-8. HP-43 transient short-term RPV code-to-data level comparison

The code-to-data comparison of CNV pressure presented in Figure B-9 shows reasonable-to-excellent agreement with a {{ }}2(a),(b),(c),ECI by NRELAP5.

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{{

}}2(a),(b),(c),ECI

Figure B-9. HP-43 transient short-term CNV pressure code-to-data comparison

The code-to-data comparison of the spurious RVV orifice mass flow rate presented in Figure B-10 shows less than reasonable agreement for {{ }}2(a),(b),(c),ECI of the transient, with reasonable predictions thereafter. However, test data in Figure B-6 (PZR pressure) and Figure B-9 (CNV pressure) show {{

}}2(a),(b),(c),ECI

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{{

}}2(a),(b),(c),ECI

Figure B-10. HP-43 transient short-term spurious RVV orifice mass flow rate code-to-data comparison

The code-to-data comparison of CNV level presented in Figure B-11 shows reasonable- to-excellent agreement with a slight under-prediction by NRELAP5 {{ }}2(a),(b),(c),ECI

The trends in CNV pressure and level presented in Figure B-9 and Figure B-11 appear to indicate {{

}}2(a),(b),(c),ECI

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{{

}}2(a),(b),(c),ECI

Figure B-11. HP-43 transient short-term CNV level code-to-data comparison

B.7.4 Applicability of NRELAP5 LOCA EM to IORV Analysis

Section 8.0 of this report demonstrates the adequacy of the NRELAP5 code for the analysis of design-basis LOCAs in the NPM by use of closure model and correlation reviews, and assessments against relevant experimental data. Because there are no significant differences in physics phenomena between the LOCA and IORV events, the code assessments performed in support of the LOCA EM are also applicable for the IORV EM.

The dominant code models and correlations for the LOCA PIRT shown in Table 8-1 are also applicable for the IORV EM. The range of NPM key process parameters for LOCA shown in Table 8-2 are also applicable for IORV analysis, such that the LOCA parameter ranges are identical to or envelope the IORV ranges.

B.8 IORV Analysis Results

The IORV analysis example results and sensitivity studies are presented in this section with the objective of supporting the development of the IORV EM, and demonstrating its successful application for the evaluation of the AOO acceptance criteria for postulated IORV events.

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• Section B.8.1 briefly presents a typical sequence of events for both the inadvertent RVV and RRV opening cases. The RVV event sequence assumes no loss of electrical power, while the RRV event sequence assumes a loss of normal AC and DC power at event initiation. The event sequences show how the availability of normal electric power affects the ECCS timing. The addition of ECCS actuation on low RCS pressure would result in earlier actuation of the remaining ECCS valves for an inadvertent opening of an RVV with power available. For an inadvertent opening of an RRV with power available, the remaining ECCS valves would still open on the high CNV level actuation signal. IORV cases with a loss of all power are unaffected by the addition of the low RCS pressure signal (remaining ECCS valves open at IAB threshold). • Section B.8.2 presents the matrix of initial condition biases applied to the IORV calculations, the results of which are shown in Sections B.8.3 through B.8.5, and support development of methodology guidance for biasing the initial conditions for IORV analysis cases. • Section B.8.3 presents analysis results for inadvertent RVV opening for the 17 initial condition biases with no loss of electrical power. • Section B.8.4 presents analysis results for inadvertent RRV opening for the 17 initial condition biases with no loss of electrical power. • Section B.8.5 presents analysis results for inadvertent RSV opening for the 17 initial condition biases with no loss of electrical power. • Section B.8.6 presents analysis results for sensitivity cases on model parameters, including: − Fuel rod gap conductance − Axial power shape − ECCS valve sizing − ECCS valve opening rate − DHRS availability/credit − Assumed single active failures • Section B.8.7 presents analysis results from applying the single active failure assumptions listed in Section B.6.2 to the limiting RVV and RRV base cases. • Section B.8.8 presents analysis results from applying the three electrical power availability assumptions listed in Section B.6.1 to the limiting RVV and RRV base cases. • Section B.8.9 presents plots of parameters of interest for the typical RVV opening and RRV opening cases presented in Sections B.8.1 and B.8.2.

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B.8.1 IORV Event Progression in the NuScale Power Module

B.8.1.1 RVV Sequence of Events

The sequence of events for a typical inadvertent RVV opening without loss of electrical power is shown in Table B-6. The time of MCHFR occurs very early in the transient, before the control rods are fully inserted from the reactor trip. Peak steam generator and peak containment pressure occurs soon after the reactor trip and containment isolation, within the first 50 seconds. The remaining ECCS valves open much later, when the ECCS system actuates on high containment level. Natural circulation flow is established back to the RCS after containment and RCS pressures equalize across the RRVs. Minimum water level above the core occurs as the RPV and containment water levels equalize.

Table B-6. Sequence of Events for RVV opening without loss of normal AC or DC power

Event Time [s] Transient initiation due to inadvertent opening of an RVV. Peak RPV pressure 0 occurs at time zero.

High containment pressure RTS analytical limit reached 0.27

Transient minimum critical heat flux ratio occurs

The RVV blowdown into containment from the pressurizer causes rapid depressurization of the pressurizer steam space, a reduction in RCS loop flow as coolant surges into the pressurizer, an increase in nucleate boiling in the core, and a corresponding increase in the cladding heat flux to the coolant. These factors 0.34 combine to cause an immediate reduction in CHF. Following the occurrence of transient MCHFR, a temporary increase in core inlet flow is observed, caused by an increased density gradient due to voiding in the riser. The increase in flow, coupled with the reactor scram, restores CHFR margin which is maintained for the remainder of the transient.

RTS actuation signal (after 2-second delay) 2.27

Containment Isolation 4.27

Control rods fully inserted following reactor scram 4.47

Peak steam generator pressure is reached

An increase in steam generator pressure results from secondary system isolation following the reactor trip and containment isolation. Heat transfer from the RCS to 35 the secondary side inventory remaining in the steam generator decreases as the RCS pressure drops and the SG pressure increases. Consequently the IORV event scenario is not limiting for SG peak pressure.

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Event Time [s]

Peak containment pressure is reached

Containment pressure increases as coolant is transferred from the RCS into the CNV through the open RVV. Decreasing valve flow and condensation inside 49 containment cause pressure to decrease after reaching its peak value. (Note that the limiting maximum CNV pressure and temperature for AOOs is determined via separate calculations.)

ECCS actuation on high containment level1

ECCS actuates on the high containment level setpoint. At this point in the transient 2018 the pressure difference between the RPV and CNV is below the IAB threshold pressure, and the remaining ECCS valves open immediately.

Natural-circulation ECCS flow is established

The pressure drop across the RRVs equalizes, allowing liquid coolant to flow from 2445 containment back into the RPV downcomer. This establishes a two-phase natural circulation loop through the ECCS, which transfers decay heat to the reactor pool. Pressure and temperature inside the RPV and CNV continue to decrease.

Minimum collapsed liquid level above the core

Collapsed liquid level above the core active fuel continues decreasing until reaching 2445-2565 an equilibrium level at approximately 10 ft. This level is maintained for the remainder of the transient.

End of calculation.

The transient is terminated 120 seconds after natural circulation ECCS flow was established at 2445 sec. During this time stable ECCS cooling continues while RCS 2565 pressure and temperature decrease. The analysis is terminated with the NPM in a stable safe condition with RPV liquid level maintained above the active core during the entire transient.

Note: 1. An ECCS actuation signal on low RCS pressure was added to the NuScale design but is not included in the analyses in this report. This signal could actuate ECCS earlier in some valve opening events but does not significantly impact accident progression or results.

B.8.1.2 RRV Sequence of Events

The sequence of events for a typical inadvertent RRV opening with loss of normal AC and DC power is shown in Table B-7. The time of MCHFR occurs very early in the transient, before the control rods are fully inserted from the reactor trip. The remaining ECCS valves open as soon as the IAB threshold pressure is reached. Peak containment pressure occurs soon after the ECCS valves open. Natural circulation flow is established

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back to the RCS after containment and RCS pressures equalize across the RRVs. Minimum water level above the core occurs as the RPV and containment water levels equalize.

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Table B-7. Sequence of Events for RRV opening with loss of normal AC and DC power

Event Time [s]

Transient initiation due to inadvertent opening of an RRV. Peak RPV pressure occurs at time zero. 0

The control rods are fully inserted by 2.3 seconds into the transient.

Assumed loss of AC and DC power An assumed loss of AC and DC power at time zero results in a loss of feedwater 0 flow and immediate reactor scram.

Transient minimum critical heat flux ratio occurs

The RRV blowdown into containment via the downcomer leg of the primary system flow loop causes an immediate reduction in core inlet flow and a reduction in CHF. 0.5 Following the occurrence of transient MCHFR, a temporary increase in core inlet flow is observed, caused by an increased density gradient due to voiding in the riser. The increase in flow, coupled with the reactor scram, restores CHFR margin which is maintained for the remainder of the transient.

Control rods fully inserted following reactor scram 2.3

ECCS actuation at IAB release pressure

The assumed loss of DC power at event initiation would normally allow all ECCS valves to immediately open, however the IAB prevents this actuation as long as the 50 pressure difference between the RPV and CNV is greater than the IAB threshold pressure setpoint. By 50 seconds, the differential pressure has decreased below the IAB release pressure and the remaining ECCS valves open.

Peak containment pressure is reached

Containment pressure increases as coolant is transferred from the RCS into the CNV through the ECCS valves. Decreasing valve flow and condensation inside 64 containment cause pressure to decrease after reaching its peak value. (Note that the limiting maximum CNV pressure and temperature for AOOs is determined via separate calculations.)

Natural-circulation ECCS flow is established

The pressure drop across the RRVs equalizes, allowing liquid coolant to flow from 483 containment back into the RPV downcomer. This establishes a two-phase natural circulation loop through the ECCS, which transfers decay heat to the reactor pool. Pressure and temperature inside the RPV and CNV continue to decrease.

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Event Time [s]

Peak steam generator pressure is reached

An increase in steam generator pressure results from secondary system isolation following the assumed loss of AC and DC power at transient initiation. Heat transfer 490 from the RCS to the feedwater condensate remaining in the steam generator is limited, because of falling RCS saturation temperature associated with decreasing RCS pressure. Consequently the IORV event scenario is not limiting for SG peak pressure.

Minimum collapsed liquid level above the core

Collapsed liquid level above the core active fuel continues decreasing until reaching 630 an equilibrium level at approximately 10 ft. This level is maintained for the remainder of the transient.

End of calculation.

The transient is terminated 30 minutes (1800 sec) after natural circulation ECCS flow was established at 483 sec. During this time stable ECCS cooling continues 2284 while RCS pressure and temperature decrease. The analysis is terminated with the NPM in a stable safe condition with RPV liquid level maintained above the active core during the entire transient.

B.8.2 Initial Conditions

The IORV analyses and sensitivity studies thoroughly investigate the scenario variations for each RPV valve (RVV, RRV, or RSV) that is assumed to inadvertently open. Each valve opening scenario is analyzed with electrical power available. Table B-8 shows the seventeen primary initial condition sensitivities that are conducted for each of the valves. Bias condition number 17 represents the nominal module response.

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Table B-8. IORV analysis initial conditions

{{

}}2(a),(c)

RCS Initial Temperature: Analyses are performed with a highest operationally allowed RCS average temperature which places the RCS closer to saturated conditions at the time of transient initiation. Analyses are also performed at the coldest operationally allowed RCS average temperature which reduces level swell following inadvertent valve opening and lengthens the time before two-phase choking conditions exist at the open valve.

RCS Initial Flow: Analyses are performed for both minimum and maximum initial RCS flow rates. The low RCS flow condition minimizes the CHFR at transient initiation. The high flow initial condition reduces the temperature difference across the core, thereby raising the core inlet temperature when the RCS average temperature is held fixed. The values for minimum and maximum RCS initial flow are not affected by the addition of the riser diverse flow holes.

RCS Initial Pressure: Analyses are performed for both low and high RCS initial pressure conditions. The high initial pressure maximizes the initial RCS energy and results in a higher flow rate on initial RPV opening. A low initial pressure places the RCS closer to saturation at transient initiation, which increases void generation and swell while the core power is still high.

Pressurizer Initial Level: Analyses are performed to vary the initial pressurizer level between the programmed operational setpoint plus or minus instrument error. A high

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initial level reduces the initial steam volume in the pressurizer, resulting in a faster RCS depressurization following inadvertent RSV or RVV opening. However a high initial level results in earlier two-phase choking due to level swell, reducing flow through the open RVV or RSV. A low initial pressurizer level delays the onset of two-phase choking for an inadvertent RVV or RSV opening, and also reduces the available coolant to maintain level above the core later in the transient.

B.8.3 RVV Inadvertent Opening

The inadvertent opening of an RVV results in a steam space blowdown from the pressurizer to containment. This causes a pressure decrease in the primary system and a subsequent corresponding pressure increase in the containment. The high CNV pressure analytical limit is reached less than a second into the event, followed by reactor trip after a 2 second delay. The MCHFR occurs during the time between the high CNV pressure analytical limit and the reactor trip actuation. During this time the reactor power and primary coolant temperature are still relatively high, and primary flow is decreasing as coolant is drawn upward into the pressurizer during the blowdown.

Primary coolant continues to flow through the inadvertently opened RVV into containment until the level inside containment reaches the setpoint for ECCS actuation. At that time the pressure difference is below the IAB threshold pressure, thus the remainder of the RVVs and the RRVs open immediately following the ECCS actuation signal on high containment level. After the other ECCS valves open, the RPV water level drops more quickly to equalize with the containment level. The collapsed liquid level above the active fuel equalizes at about 10 ft. Table B-9 shows parameters of interest and timing for all RVV with power sensitivity cases. The initial conditions corresponding to the input bias number are shown in Table B-8.

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Table B-9. Results for RVV cases with power available

{{

}}2(a),(c)

It can be seen from these power-available RVV results in Table B-9 that cases with high initial RCS average temperature and low initial RC flow typically generated the lowest MCHFR. When these two biases are in play, the MCHFR results in this set of analysis cases are not particularly sensitive to variations in initial pressurizer pressure or initial pressurizer level.

B.8.4 RRV Inadvertent Opening

The inadvertent opening of an RRV results in a liquid space blowdown from the RPV downcomer to the containment. Similar to the RVV event, the high CNV pressure analytical limit is reached less than a second into the event, followed by reactor trip after a 2 second delay. MCHFR occurs before the high containment pressure analytical limit is reached. The overall RRV transient is similar to the RVV. However, the liquid-space discharge results in a slower depressurization, accompanied by a greater decrease in core inlet flow as coolant discharges from the downcomer region into containment.

The liquid-space discharge generates an ECCS actuation signal on high containment level that occurs earlier than for the RVV transient. After the remaining ECCS valves open, the RRV scenario and the RVV scenario follow similar trends for fluid conditions and heat transfer. Table B-10 shows parameters of interest for all RRV cases with electric power available.

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Table B-10. Results for RRV cases with power available

{{

}}2(a),(c)

For these power-available RRV results in Table B-10, cases with high initial RCS average temperature, low initial RC flow, and high initial RC pressure typically generated the lowest MCHFR. Cases with low initial pressure instead of high pressure generated marginally higher MCHFR values. The next lowest MCHFR results occur for cases with low initial RC average temperature, low initial RC flow, and high initial RC pressure. None of the MCHFR results for the cases analyzed in this set are particularly sensitive to initial pressurizer level.

The RRV case with high initial RC average temperature, low initial RC flow, low initial RC pressure, and low initial pressurizer level typically resulted in the lowest minimum water level above the core (7.7 ft) of all IORV cases analyzed.

B.8.5 RSV Inadvertent Opening

B.8.5.1 Electric Power Available

Similar to the RVV and RRV cases, the RSV with power cases result in a reactor trip when the high containment pressure analytical limit is reached. However, this occurs later due to the smaller size of the RSV compared to the ECCS valves. The overall response is similar to the RVV event. However the rate of depressurization is less, and ECCS actuation occurs later than in either the RRV or RVV scenarios. The slower event progression and smaller mass flow rate into containment results in higher (less limiting)

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CHF ratios than for the RVV scenarios. Table B-11 shows a summary of results for the RSV cases with power available.

Table B-11. Results for RSV cases with power available

{{

}}2(a),(c)

As expected the RSV power-available results in Table B-11 generally parallel the RVV power-available results in Table B-9, except that the MCHFR values are consistently higher. The MCHFR timing still occurs within the first second of transient initiation even though the reactor trip occurs much later (12-14 sec) for the RSV power-available cases.

B.8.6 Sensitivity Analyses for Model Parameters

B.8.6.1 Fuel Rod Gap Conductance

Table B-12 shows the results of analyses performed to evaluate the effect of applying the maximum fuel gap conductance instead of the minimum gap conductance used in all other cases. The RRV case shows that a minimum gap conductance is slightly more conservative, while the RVV case shows that a maximum gap conductance is slightly more conservative. The overall impact of gap conductance on MCHFR is minor for both cases, especially considering the magnitude difference between the minimum and maximum gap conductance values.

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Table B-12. Gap conductance results

{{

}}2(a),(c)

B.8.6.2 Axial Power Shape

Table B-13 shows the results of sensitivity analyses performed to evaluate the effect of core axial power shape (bottom, middle, and top-peaked) on the acceptance criteria related results. The results indicate that the middle-peaked power shape (used in all other cases) is limiting in terms of MCHFR for both the RVV and RRV events. This is consistent with expectations because the middle-peaked shape results in the highest axial peaking factor. Despite the influence on MCHFR, the axial power shape has little effect on the overall transient progression.

Table B-13. Axial power shape results

{{

}}2(a),(c)

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B.8.6.3 ECCS Valve Sizing

It is generally understood that maximizing the RVV and RRV flow capacities results in faster RCS depressurization and thus minimizes CHFR. However, sensitivity analyses are performed on the limiting case scenarios with minimum ECCS valve sizing to confirm this assumption. The results shown in Table B-14 confirm that the maximum ECCS valve size is more conservative for MCHFR.

Table B-14. ECCS valve capacity results

{{

}}2(a),(c)

B.8.6.4 ECCS Valve Opening Stroke Time

It is expected that faster ECCS valve opening times will result in a higher total mass flow through the inadvertently opened valve with faster RPV depressurization, which should also result in a lower MCHFR. Sensitivity analyses are performed using longer ECCS valve opening times to confirm this assumption. The results shown in Table B-15 confirm that the faster ECCS valve opening time is more conservative for MCHFR.

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Table B-15. ECCS valve stroke time results

{{

}}2(a),(c)

B.8.6.5 DHRS Operation

DHRS operation is not credited in the IORV analyses. Table B-16 shows the results of sensitivity analyses performed to confirm that normal DHRS operation does not make the results worse for the limiting RVV and RRV case scenarios. The DHRS valves open after the time of MCHFR and thus the system has no effect on MCHFR.

Table B-16. DHRS operation results

{{

}}2(a),(c)

B.8.7 Single Active Failures

The maximum flow capacity through the ECCS valves is achieved when no single active failures are applied. This maximizes RCS depressurization which is limiting for MCHFR. Also, because the time of MCHFR occurs very early in the transient, the assumed failures should have no effect on MCHFR. Table B-17 shows the results of sensitivity

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cases performed by applying the single active failures from Section B.6.2 to the RVV and RRV limiting case scenarios to confirm that the base case results remain limiting.

Table B-17. Single failure results

{{

}}2(a),(c)

B.8.8 Electric Power Availability

The three electric power scenarios identified in Section B.6.1 are evaluated for the limiting cases to confirm that the most limiting scenario has been identified. The results in Table B-18 show MCHFR has negligible sensitivity to a loss of power. The transient timing is early enough such that any variation in thermal-hydraulic conditions caused by a loss of power has negligible impact on core conditions before MCHFR occurs. The loss of AC power scenario is similar to that when all power is available. A loss of DC power results in earlier ECCS cooling since the valves open as soon as the IAB release pressure is reached.

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Table B-18. Electric power availability results

{{

}}2(a),(c)

B.8.9 Plots of Parameters of Interest

B.8.9.1 Inadvertent RVV Opening Plots

The following figures are generated from case #1 in Table B-9, for the inadvertent opening of an RVV with normal AC and DC electrical power remaining available during the event. The sequence of events shown in Table B-6 also accompanies these figures.

The transient begins with an inadvertent opening of a single RVV which initiates flow from the pressurizer to the containment. The initial flow peaks at approximately 900 lbm/s (Figure B-12) and trends downward with oscillations over the next 50 seconds. The steam and two-phase flow from the pressurizer into containment causes the RPV pressure to fall and the CNV pressure to increase (Figure B-14). The RPV and CNV pressures have equalized by approximately 50 seconds and they trend together for the remainder of the analysis (Figure B-15). AC and DC power is assumed maintained for the duration of the event, therefore the opening of the remaining ECCS valves is delayed until shortly after 2000 seconds when the ECCS actuation signal on high CNV level has occurred and the IAB pressure threshold has been reached (Figure B-28 and Figure B-29).

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Figure B-12. Inadvertently opened RVV flow (short term)

Figure B-13. Inadvertently opened RVV flow (long term)

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Figure B-14. RPV and CNV pressure (short term) for inadvertent RVV opening

Figure B-15. RPV and CNV pressure (long term) for inadvertent RVV opening

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A reactor trip occurs on high CNV pressure (after a 2-seond delay) and the control rods are fully inserted by 4.47 seconds (Figure B-25). The power rapidly decreases to approximately 100 MWt and then increases back up to approximately 140 MWt in response to net reactivity feedback (Figure B-26), before control rod insertion takes effect. The core power does not increase after the control rods are inserted.

The flow from the pressurizer through the open RVV initially causes core flow to increase as the RCS flow surges towards the pressurizer as shown in Figure B-16. The RCS flow then decreases, and increasingly damped surges continue until the flow stabilizes after approximately 100 seconds (Figure B-17).

Figure B-16. RCS flow (short term) for inadvertent RVV opening

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Figure B-17. RCS flow (long term) for inadvertent RVV opening

The RCS average temperature and the core outlet temperature initially decrease due to the initial increase in core flow combined with the cooling effects of enhanced core nucleate boiling due to the drop in RPV pressure from the open RVV (Figure B-18). Temperatures rise and then decrease slightly at approximately 30 seconds in response to RCS flow surges. The RCS temperatures converge and follow the saturation line as the RPV pressure falls (Figure B-19).

The calculated CHFR reaches a minimum very early in the event, at approximately 0.34 seconds as shown in Figure B-20 and Figure B-22. For the inadvertent RVV scenario the minimum CHFR condition is influenced most by the increase in core heat flux to the coolant as the depressurization and void generation initially enhances heat transfer. Following reactor trip the CHFR increases dramatically and stays high for the remainder of the analysis as shown in Figure B-21.

Following the inadvertent RVV opening the RPV water level gradually decreases as shown in Figure B-23 and Figure B-24. The RPV and CNV levels approach equilibrium after the opening of the ECCS valves at approximately 2000 seconds. The minimum water level above the active fuel is approximately 10 feet. Since the core remains covered and MCHFR remains above the 95/95 limit, the AOO fuel centerline temperature limits are not challenged. The volume-averaged fuel temperatures decrease following the reactor trip and continue to decrease over the duration of the analysis (Figure B-30 and Figure B-31).

Figure B-27 shows the steam pressure response for SG-1. The steam pressure quickly peaks at approximately 815 psia following reactor/turbine trip and then gradually

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decreases until ECCS actuation after which the pressure remains stable at approximately 350 psia. The IORV events do not challenge the SG pressure limits and are bounded by other AOO events in this regard.

Figure B-18. RCS temperatures (short term) for inadvertent RVV opening

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Figure B-19. RCS temperatures (long term) for inadvertent RVV opening

Figure B-20. CHFR (short term) for inadvertent RVV opening

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Figure B-21. CHFR (long term) for inadvertent RVV opening

Figure B-22. Transient MCHFR for inadvertent RVV opening

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Figure B-23. RPV and CNV level (short term) for inadvertent RVV opening

Figure B-24. RPV and CNV level (long term) for inadvertent RVV opening

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Figure B-25. Reactor Power (short term) for inadvertent RVV opening

Figure B-26. Net reactivity (short term) for inadvertent RVV opening

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Figure B-27. SG-1 pressure for inadvertent RVV opening

Figure B-28. ECCS (non-inadvertently opened) RVV flow

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Figure B-29. ECCS (non-inadvertently opened) RRV flow

Figure B-30. Fuel temperature (°F) for inadvertent RVV opening (short term)

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Figure B-31. Fuel temperature (°F) for inadvertent RVV opening (long term)

B.8.9.2 Inadvertent RRV Opening Plots

The following figures are generated from case #1 in Table B-9 for the inadvertent opening of an RRV, however a loss of normal AC and DC electrical power at event initiation is also applied to demonstrate the impact on event sequence. The sequence of events shown in Table B-7 also accompanies these figures.

The transient begins with an inadvertent opening of a single RRV which initiates flow from the RPV downcomer to the containment. The initial flow peaks at approximately 540 lbm/s (Figure B-32) and gradually decreases over the next 50 seconds. The coolant flow from the RPV into containment causes the RPV pressure to fall and the CNV pressure to increase (Figure B-34 and Figure B-35). The assumed loss of DC power at event initiation would normally allow all ECCS valves to immediately open, however the IAB prevents this actuation as long as the pressure difference between the RPV and CNV is greater than the IAB pressure setpoint. The IAB threshold pressure is reached at 50 seconds and the remaining ECCS valves open (Figure B-48).

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Figure B-32. Inadvertently opened RRV flow (short term)

Figure B-33. Inadvertently opened RRV flow (long term)

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Figure B-34. RPV and CNV pressure (short term) for inadvertent RRV opening

Figure B-35. RPV and CNV pressure (long term) for inadvertent RRV opening

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An immediate reactor trip occurs due to the assumed loss of normal DC power, and the control rods are fully inserted by 2.3 seconds (Figure B-45). The total reactivity remains negative after the reactor trip (Figure B-46) and the core power does not increase following the reactor trip.

The flow loss from the RPV downcomer through the open RRV causes an immediate decrease in RCS flow as shown in Figure B-36. After the initial decrease the RCS flow temporarily recovers due to an increasing loop density gradient caused by increased voiding in the riser above the core. The recovery is short-lived and the flow begins to decrease again after approximately 8 seconds.

Figure B-36. RCS flow (short term) for inadvertent RRV opening

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Figure B-37. RCS flow (long term) for inadvertent RRV opening

The RCS average temperature and the core outlet temperature initially increase due to the reduction in core flow and the isolation of SG feedwater and steam flow following the reactor trip (Figure B-38). Temperatures begin to decrease after approximately 5 seconds due to the core power decrease following reactor trip (Figure B-45). Following ECCS actuation the RCS temperatures converge and follow the saturation line as the RPV pressure falls (Figure B-39).

The calculated CHFR reaches a minimum very early in the event, at approximately 0.5 seconds as shown in Figure B-40 and Figure B-42. For the inadvertent RRV scenario the minimum CHFR condition is influenced most by the reduction in core inlet flow. Following reactor trip the CHFR increases dramatically and stays high for the remainder of the analysis as shown in Figure B-41.

Following the opening of the ECCS valves at 50 seconds, the RPV water level decreases and reaches equilibrium with the CNV water level as shown in Figure B-43 and Figure B-44. The minimum water level above the active fuel is approximately 10 feet. Since the core remains covered and MCHFR remains above the 95/95 limit, the AOO fuel centerline temperature limits are not challenged. The volume-averaged fuel temperatures decrease following the reactor trip and continue to decrease over the duration of the analysis (Figure B-49 and Figure B-50).

Figure B-47 shows the steam pressure response for SG-1. The IORV events do not challenge the SG pressure limits and are bounded by other AOO events in this regard.

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Figure B-38. RCS temperatures (short term) for inadvertent RRV opening

Figure B-39. RCS temperatures (long term) for inadvertent RRV opening

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Figure B-40. CHFR (short term) for inadvertent RRV opening

Figure B-41. CHFR (long term) for inadvertent RRV opening

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Figure B-42. Transient MCHFR for inadvertent RRV opening

Figure B-43. RPV and CNV level (short term) for inadvertent RRV opening

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Figure B-44. RPV and CNV level (long term) for inadvertent RRV opening

Figure B-45. Reactor Power (short term) for inadvertent RRV opening

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Figure B-46. Net reactivity (short term) for inadvertent RRV opening

Figure B-47. SG-1 pressure for inadvertent RRV opening

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Figure B-48. ECCS (non-inadvertently opened) RRV flow

Figure B-49. Fuel temperature (°F) for inadvertent RRV opening (short term)

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Figure B-50. Fuel temperature (°F) for inadvertent RRV opening (long term)

B.9 IORV EM Conclusions

It has been shown that there are no differences in physical phenomena between the LOCA and IORV events, and although the initiating events are different, the governing thermal hydraulic and core physics mechanisms are the same. Therefore the high- ranked phenomena from the LOCA PIRT also apply to the IORV event scenarios. The PIRT event phases of initial blowdown (1a) and ECCS actuation (1b) are also identical for LOCA and IORV.

The NRELAP5 SET and IET assessments conducted for the LOCA EM also support NRELAP5 for use for IORV analysis. The NIST-1 HP-43 assessment for inadvertent RVV opening has been added to the original validation base provided by the HP-09 RVV opening assessment.

The NRELAP5 CHF correlation has been shown to be applicable for IORV analysis, using the KATHY CHF test data to derive a 95/95 CHF limit appropriate for use in AOO analysis.

Sample NRELAP5 calculation results for both inadvertent RVV opening and inadvertent RRV opening show results that are consistent with expectations from both governing thermal-hydraulic and core kinetics phenomena, and from the LOCA EM.

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It is therefore concluded that the extension of the NRELAP5 LOCA Evaluation Model (with minor modifications) to evaluate IORV AOO event scenarios is both feasible and appropriate.

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Appendix C. Spurious Reactor Recirculation Valve Opening Integral Effects Test

C.1 Purpose

The HP-49 test was performed at the NIST-1 facility and was used to assess the capability of NRELAP5 to predict the integral response of the NIST-1 facility for a spurious reactor recirculation valve (RRV) opening inside containment. The reactor recirculation line (RRL) and RRV connect the downcomer side of the RPV to the CNV.

C.2 Facility Description

The NIST-1 facility is described in Section 7.5.1. The entire NIST-1 facility except for the CVCS, PZR Spray, and DHRS was used for this IET, including:

• the SG was active to remove heat from the primary side and drive natural circulation in conjunction with the electrically heated core during the steady state period • {{

}}2(a),(b),(c) • the CPV was filled to accept rejected heat from the HTP

C.3 Phenomenon Addressed

The HP-49 test is an IET modeling a spurious RRV opening into containment. The pertinent phenomena addressed by this test are:

• {{

}}2(a),(b),(c),ECI

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C.4 Experimental Procedure

The experiment test procedure is consistent with the LOCA test procedure described in Section 7.5.1.6. When the CNV pressure reached the specified CNV transient initiation pressure, the spurious RRV was opened, initiating the transient.

Within the NIST-1 facility, the ECCS actuation occurs when the compensated level in the RPV downcomer reads lower than a specified value. Once this occurs, open signals are sent to the remaining RRV and the RVVs. The opening of the ECCS valves causes a large amount of mass and energy transfer to occur between the RPV and the CNV over a short period of time. The CNV pressurization and heat-up occurs rapidly, followed by a long depressurization and cooldown profile. Test data was recorded for an extended period of time, into the long-term cooling phase.

C.5 Special Analysis Techniques

The RRV discharge line orifice has a length of approximately {{ . }}2(a),(b),(c),ECI and an ID of approximately {{ }}2(a),(b),(c),ECI Thus, the orifice has an L/D ratio roughly equal to {{ }}2(a),(b),(c),ECI Analysis indicates that an NRELAP5 discharge coefficient near {{ }}2(a),(b),(c),ECI produces reasonable agreement with the spurious RRV flow rate inferred from test data, however, the literature determined value of {{ }}2(a),(b),(c),ECI is used for the base case assessment.

The {{

}}2(a),(c)

C.6 Assessment Results (HP-49)

The NRELAP5 transient model is designed to simulate initial test conditions and includes logic that follows facility controls and test procedures. For this experiment, the spurious mass flow rate was not measured. The calculated spurious flow rate is reasonable because the differential pressure across the spurious RRV line orifice (Figure C-1), the RPV level response (Figure C-4), the CNV level response (Figure C-5), the RPV pressure response (Figure C-8), and the CNV pressure response (Figure C-6) are all in reasonable agreement for the pre-ECCS opening period of the transient.

The NIST-1 v-cone flowmeter (measuring primary loop flowrate) is designed for positive single-phase liquid conditions. During the HP-49 test, two-phase conditions occur at the location of the v-cone meter. As shown in Figure C-2 NRLEAP5 captures the RPV primary-flow coast-down period after transient initiation with reasonable accuracy. Note that after {{

}}2(a),(b),(c),ECI

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The pressurizer level is compared in Figure C-3. The comparisons show reasonable agreement. NRELAP5 predicts complete draining of the pressurizer at about {{ }}2(a),(b),(c),ECI

NRELAP5 provides reasonable agreement for level response in the RPV and CNV as shown in Figure C-4 and Figure C-5. The CNV peak pressure and pressure response are also predicted with reasonable agreement to data as shown in Figure C-6 and Figure C-7. The timing of ECCS actuation is predicted with reasonable agreement to the test data. Primary pressure response is predicted with reasonable agreement (Figure C-8).

{{

2(a),(b),(c),ECI }}

Figure C-1. NIST-1 HP-49 spurious orifice differential pressure

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{{

2(a),(b),(c),ECI }} Figure C-2. NIST-1 HP-49 primary mass flow rate

{{

}}2(a),(b),(c),ECI

Figure C-3. NIST-1 HP-49 pressurizer level comparison

TR-0516-49422-NP-A Rev. 2

{{

2(a),(b),(c),ECI }} Figure C-4. NIST-1 HP-49 reactor pressure vessel level comparison

{{

}}2(a),(b),(c),ECI

Figure C-5. NIST-1 HP-49 containment vessel level comparison

TR-0516-49422-NP-A Rev. 2

{{

}}2(a),(b),(c),ECI

Figure C-6. NIST-1 HP-49 containment vessel peak pressure comparison

{{

}}2(a),(b),(c),ECI

Figure C-7. NIST-1 HP-49 containment vessel pressure comparison

TR-0516-49422-NP-A Rev. 2

{{

}}2(a),(b),(c),ECI

Figure C-8. NIST-1 HP-49 primary pressure comparison

TR-0516-49422-NP-A Rev 2

Section C

TR-0516-49422-P-A Rev. 2

Requests for Additional Information: Loss‐of‐Coolant Accident Evaluation Model

Changes from the RAIs listed below have been incorporated into the body of this report, and the full contents of all listed RAIs and their responses are part of the approved version of this topical report, and required for the use of this methodology. The last entry in this table includes changes associated with NuScale letter as indicated, and is not associated with an RAI response. Set eRAI Question NuScale Letter Non‐Proprietary Proprietary Topical Report Number Number Number Number Accession Accession Number Sections Affected Number 8776 8776 15.06.05‐2 RAIO‐1017‐56660 ML17291B321 ML17291B322 N/A 15.06.05‐3 15.06.05‐4 15.06.05‐5 15.06.05‐6 8777 8777 15.06.05‐1 RAIO‐0917‐56266 ML17270A310 ML17270A311 Section 7.5.1 ‐ Containment wall heat transfer 8985 8985 15.06.06‐1 RAIO‐1117‐57045 ML17310B505 ML17310B506 N/A 8990 8990 15.06.05‐7 RAIO‐1117‐57291 ML17324B392 ML17324B393 N/A 8990 8990 15.06.05‐7 RAIO‐0819‐66823 ML19240C658 ML19240C659 Sections 6.8, 8.2 and 8.3 ‐ (supplemental) Containment Wall Heat Transfer and Condensation modeling 9065 9065 15.06.05‐8 RAIO‐1217‐57788 ML17353A951 ML17353A952 N/A 9085 9085 15.06.05‐12 RAIO‐0918‐61956 ML18271A158 ML18271A160 Sections 5, 9 and Appendix A ‐ Clarified Steam Generator modeling 9126 9126 15.06.05‐9 RAIO‐0118‐58482 ML18031B319 ML18031B320 N/A 9149 9149 15.06.05‐10 RAIO‐0118‐58451 ML18030B256 ML18030B258 N/A 9190 9190 15.06.05‐11 RAIO‐0218‐58565 ML18038B602 N/A N/A 9208 9208 15.06.05‐14 RAIO‐0819‐66821 ML19240C644 ML19240C665 Sections 6.8, 8.2 and 8.3 ‐ (supplemental) Containment Wall Heat

Loss-of-Coolant Accident Evaluation Model

TR-0516-49422-P-A Rev. 2

Set eRAI Question NuScale Letter Non‐Proprietary Proprietary Topical Report Number Number Number Number Accession Accession Number Sections Affected Number Transfer and Condensation modeling 9208 9208 15.06.05‐14 RAIO‐0219‐64682 ML19058A864 ML19058A865 Sections 8.3.2 and 8.3.4 ‐ 15.06.05‐15 Scaling and Scaling 15.06.05‐16 Effects Distortion clarified 9390 9390 15.06.05‐ RAIO‐0219‐64680 ML19058A867 ML19058A868 Sections 8.3.2 and 8.3.4 ‐ 18(d) Scaling and Scaling 15.06.05‐19 Effects Distortion clarified 9390 9390 15.06.05‐18 RAIO‐0119‐64307 ML19029B559 ML19029B560 N/A (a), (b), (c) 9390 9390 15.06.05‐18 RAIO‐0919‐66910 ML19252A434 ML19252A437 N/A (supplemental) 9475 9475 15.06.05‐17 RAIO‐0918‐61890 ML18271A167 ML18271A169 Sections 5, 9 and Appendix A ‐ Clarified Steam Generator, Containment Vessel, Secondary System modeling 9476 9476 15.06.05‐13 RAIO‐0918‐61838 ML19261A400 ML18261A401 N/A 9492 9492 15.06.05‐21 RAIO‐0918‐61793 ML18260A315 ML18260A316 N/A 9494 9494 06.02.01.01.A‐ RAIO‐0119‐64081 ML19009A550 ML19009A551 N/A 17 9519 9519 15.06.05‐20 RAIO‐0818‐61403 ML18228A826 ML18228A827 Sections 7.4, 8.2 ‐ Clarified SIET Pitch to Diameter Ratio 9536 9536 15.06.06‐2 RAIO‐0918‐61859 ML18264A338 ML18264A339 Added Appendix B ‐ Evaluation Model for

Loss-of-Coolant Accident Evaluation Model

TR-0516-49422-P-A Rev. 2

Set eRAI Question NuScale Letter Non‐Proprietary Proprietary Topical Report Number Number Number Number Accession Accession Number Sections Affected Number Inadvertent Opening of RPV Valves 9536 9536 15.06.06‐2 RAIO‐0819‐66717 ML19235A258 ML19235A259 Sections 2.2, 6.11, and (supplemental) 7.3 ‐ CHF correlation clarifications 9549 9549 06.02.01.01.A‐ RAIO‐0918‐61935 ML18267A365 ML18268A366 Added Appendix C ‐ 21 Spurious Reactor Recirculation Valve Opening Integral Effects Test N/A N/A N/A LO‐0620‐70500 ML20175A345 ML20175A346 Section 3.3 – added ECCS actuation description; Section 3.3.1 – added discussion of ECCS valve opening on low RPV and CNV differential pressure; Section 4.2 – added discussion of steam space and liquid space LOCA example calculations; section 5.1.1 – added riser holes to nodalization description; Section 5.1.2.3 – added discussion of addition of small diverse flow paths to riser; Table 5‐3 – added safety‐related detection purpose information to existing

Loss-of-Coolant Accident Evaluation Model

TR-0516-49422-P-A Rev. 2

Set eRAI Question NuScale Letter Non‐Proprietary Proprietary Topical Report Number Number Number Number Accession Accession Number Sections Affected Number measurement parameters and added the parameter of RCS pressure; Section B.8 – added ECCS actuation changes to IORV analysis results discussion; Section B.8.2 – added text regarding effect of adding diverse flow holes to the riser

Loss-of-Coolant Accident Evaluation Model

TR-0516-49422-NP-A Rev 2

Section D

May 27, 2020 Docket No. PROJ0769

U.S. Nuclear Regulatory Commission ATTN: Document Control Desk One White Flint North 11555 Rockville Pike Rockville, MD 20852-2738

SUBJECT: NuScale Power, LLC Submittal of “Loss-of-Coolant Accident Evaluation Model,” TR-0516-49422, Revision 2

REFERENCE: Letter from NuScale Power to NRC, “NuScale Power, LLC Submittal of ‘Loss-of-Coolant Accident Evaluation Model,’ TR-0516-49422, Revision 1,” dated November 27, 2019 (ML19331B585)

NuScale Power, LLC (NuScale) hereby submits Revision 2 of the “Loss-of-Coolant Accident Evaluation Model” (TR-0516-49422).

Enclosure 1 contains the proprietary version of the report entitled “Loss-of-Coolant Accident Evaluation Model.” NuScale requests that the proprietary version be withheld from public disclosure in accordance with the requirements of 10 CFR § 2.390. The enclosed affidavit (Enclosure 3) supports this request. Enclosure 1 has also been determined to contain Export Controlled Information. This information must be protected from disclosure per the requirements of 10 CFR § 810. Enclosure 2 contains the nonproprietary version of the report.

This letter makes no regulatory commitments and no revisions to any existing regulatory commitments.

If you have any questions, please feel free to contact Matthew Presson at 541-452-7531 or at [email protected].

Sincerely,

Zackary W. Rad Director, Regulatory Affairs NuScale Power, LLC

NuScale Power, LLC 1100 NE Circle Blvd., Suite 200 Corvallis, Oregon 97330 Office 541.360-0500 Fax 541.207.3928 www.nuscalepower.com LO-0520-70288 Page 2 of 2 05/27/2020

Distribution: Gregory Cranston, NRC Prosanta Chowdhury, NRC Michael Dudek, NRC Michael Snodderly, NRC Christiana Liu, NRC Christopher Brown, NRC Marieliz Johnson, NRC Rani Franovich, NRC

Enclosure 1: “Loss-of-Coolant Accident Evaluation Model,” TR-0516-49422-P, Revision 2, proprietary version Enclosure 2: “Loss-of Coolant Accident Evaluation Model,” TR-0516-49422-NP, Revision 2, nonproprietary version Enclosure 3: Affidavit of Zackary W. Rad, AF-0520-70289

NuScale Power, LLC 1100 NE Circle Blvd., Suite 200 Corvallis, Oregon 97330 Office 541.360-0500 Fax 541.207.3928 www.nuscalepower.com LO-0520-70288

Enclosure 2:

“Loss-of-Coolant Accident Evaluation Model,” TR-0516-49422-NP, Revision 2, nonproprietary version

NuScale Power, LLC 1100 NE Circle Blvd., Suite 200 Corvallis, Oregon 97330 Office 541.360-0500 Fax 541.207.3928 www.nuscalepower.com

LO-0620-70792

Enclosure 3:

Affidavit of Zackary W. Rad, AF-0620-70793

NuScale Power, LLC 1100 NE Circle Blvd., Suite 200 Corvallis, Oregon 97330 Office 541.360-0500 Fax 541.207.3928 www.nuscalepower.com

NuScale Power, LLC

AFFIDAVIT of Zackary W. Rad

I, Zackary W. Rad, state as follows:

(1) I am the Director of Regulatory Affairs of NuScale Power, LLC (NuScale), and as such, I have been specifically delegated the function of reviewing the information described in this Affidavit that NuScale seeks to have withheld from public disclosure, and am authorized to apply for its withholding on behalf of NuScale

(2) I am knowledgeable of the criteria and procedures used by NuScale in designating information as a trade secret, privileged, or as confidential commercial or financial information. This request to withhold information from public disclosure is driven by one or more of the following:

(a) The information requested to be withheld reveals distinguishing aspects of a process (or component, structure, tool, method, etc.) whose use by NuScale competitors, without a license from NuScale, would constitute a competitive economic disadvantage to NuScale. (b) The information requested to be withheld consists of supporting data, including test data, relative to a process (or component, structure, tool, method, etc.), and the application of the data secures a competitive economic advantage, as described more fully in paragraph 3 of this Affidavit. (c) Use by a competitor of the information requested to be withheld would reduce the competitor’s expenditure of resources, or improve its competitive position, in the design, manufacture, shipment, installation, assurance of quality, or licensing of a similar product. (d) The information requested to be withheld reveals cost or price information, production capabilities, budget levels, or commercial strategies of NuScale. (e) The information requested to be withheld consists of patentable ideas.

(3) Public disclosure of the information sought to be withheld is likely to cause substantial harm to NuScale’s competitive position and foreclose or reduce the availability of profit-making opportunities. The accompanying topical report reveals distinguishing aspects about NuScale’s loss-of-coolant accident evaluation model used for analysis of design-basis loss-of-coolant accidents in the NuScale power module.

NuScale has performed significant research and evaluation to develop a basis for this model and has invested significant resources, including the expenditure of a considerable sum of money.

The precise financial value of the information is difficult to quantify, but it is a key element of the design basis for a NuScale plant and, therefore, has substantial value to NuScale.

If the information were disclosed to the public, NuScale's competitors would have access to the information without purchasing the right to use it or having been required to undertake a similar expenditure of resources. Such disclosure would constitute a misappropriation of NuScale's intellectual property, and would deprive NuScale of the opportunity to exercise its competitive advantage to seek an adequate return on its investment.

(4) The information sought to be withheld is in the Enclosure 1 to the “NuScale Power, LLC Submittal of the Approved Version of NuScale Topical Report, ‘Loss-of-Coolant Accident Evaluation Model,’ TR-0516-49422, Revision 2.” The enclosure contains the designation “Proprietary" at the top of each page containing proprietary information. The information considered by NuScale to be proprietary is identified within double braces, "{{ }}" in the document.

AF-0620-70793 Page 1 of 2

(5) The basis for proposing that the information be withheld is that NuScale treats the information as a trade secret, privileged, or as confidential commercial or financial information. NuScale relies upon the exemption from disclosure set forth in the Freedom of Information Act ("FOIA"), 5 USC § 552(b)(4), as well as exemptions applicable to the NRC under 10 CFR §§ 2.390(a)(4) and 9.17(a)(4).

(6) Pursuant to the provisions set forth in 10 CFR § 2.390(b)(4), the following is provided for consideration by the Commission in determining whether the information sought to be withheld from public disclosure should be withheld:

(a) The information sought to be withheld is owned and has been held in confidence by NuScale.

(b) The information is of a sort customarily held in confidence by NuScale and, to the best of my knowledge and belief, consistently has been held in confidence by NuScale. The procedure for approval of external release of such information typically requires review by the staff manager, project manager, chief technology officer or other equivalent authority, or the manager of the cognizant marketing function (or his delegate), for technical content, competitive effect, and determination of the accuracy of the proprietary designation. Disclosures outside NuScale are limited to regulatory bodies, customers and potential customers and their agents, suppliers, licensees, and others with a legitimate need for the information, and then only in accordance with appropriate regulatory provisions or contractual agreements to maintain confidentiality.

(d) No public disclosure of the information has been made, and it is not available in public sources. All disclosures to third parties, including any required transmittals to NRC, have been made, or must be made, pursuant to regulatory provisions or contractual agreements that provide for maintenance of the information in confidence.

(e) Public disclosure of the information is likely to cause substantial harm to the competitive position of NuScale, taking into account the value of the information to NuScale, the amount of effort and money expended by NuScale in developing the information, and the difficulty others would have in acquiring or duplicating the information. The information sought to be withheld is part of NuScale's technology that provides NuScale with a competitive advantage over other firms in the industry. NuScale has invested significant human and financial capital in developing this technology and NuScale believes it would be difficult for others to duplicate the technology without access to the information sought to be withheld.

I declare under penalty of perjury that the foregoing is true and correct. Executed on July 7, 2020.

______Zackary W. Rad

AF-0620-70793 Page 2 of 2